ML20067A524

Final Submittal in Response to NRC Bulletin 88-011, 'Pressurizer Surge Line Thermal Stratification.'

ML20067A524
Person / Time
Site:	Davis Besse, Oconee, Arkansas Nuclear, Three Mile Island, Crystal River
Issue date:	12/31/1990
From:	Costa D, Gurdal R, Matthews M BABCOCK & WILCOX CO.
To:
Shared Package
ML20067A436	List: 3F0191-07, Forwards BAW-2127, Final Submittal for NRC Bulletin 88-011, 'Pressurizer Surge Line Thermal Stratification.' Revised Operating Transients Will Accommodate Normal Variations in Operations & Ensure Margin in Total Fatigue Usage Factor ML20067A524
References
BAW-2127, IEB-88-011, IEB-88-11, NUDOCS 9102070283
Download: ML20067A524 (189)

Text

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FINAL SUBMITTAL FOR NUCLEAR REGULATORY COMMISSION BULLETIN 88-11

" PRESSURIZER SURGE LINE THERMAL STRATIFICATION" O m t13gJSERVICE ee fsawmuCtean COMPANY '

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t BAW 2127 December 1990 FINAL SUBMITTAL IN RESPONSE TO NUCLEAR REGULATORY COMMISSION .

BULLETIN 88-11 4

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" PRESSURIZER SURGE LINE THERMAL STRATIFICATION" i

() Prepared for Arkansas Power.and Light Company  :

Duke Power Company l Florida Power Corporation l General Public Utilities Nuclear Toledo Edison Company Prepared by (see Section 10 for document signatures)

M.T. Matthews R.J. Gurdal:

D.E.-Costa l G.L. Weatherly B&W Nuclear Service Company P.O. Box 10935 O Lynchburg, Virginia 24506;0935-

EXECUTIVE

SUMMARY

On December 20, 1988 the Nuclear Regulatory Commission issued NRC Bulletin 88-11.

The bulletin addressed technical concerns associated with thermal stratification in the pressurizer surge line and required utilities to establish and implement a program to ensure the structural integrity of the surge line. Subsequent to issuance of the bulletin, the B&W Owners Group has developed a comprehensive program to address the requirements of the bulletin. A committee which includes representatives from each member utility has closely participated in all phases of the program. This report summarizes the B&WOG program and its results.

The surge line for each B&W designed lowered loop plant has been evaluated for the effects of thermal stratification. Davis-Besse Unit 1 (DB 1), the only B&WOG raised loop plant, is undergoing a plant specific evaluation which will be

, reported in a supplement to this report. The evaluation of the lowered loop plants involved comprehensive instrumentation of Oconee Unit 1 as the representative plant. The evaluation also involved assessment of operating

practices and procedures, collection and review of historical plant data from all six lowered loop B&W plants, and development of new design basis transient conditions for the surge line to conservatively account for thermal cycling, thermal stratification, and thermal striping. The evaluation of thermal striping incorporated the best available data to characterize this phenomenon as it may occur in the surge line.

The structural analysis for the new design conditions has shown that the lowered

' loop surge line can meet its 40 year design life. Detailed finite element analyses have been performed on the pressurizer surge nozzle, on the surge line to hot leg nozzle, and on the limiting portions (the elbows) of the pressurizer surge line piping. At all points in the serge line and the associated nozzles the cumulative fatigue usage factor remains less than one for the design life.

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CONTENTS Page EXECUTIVE

SUMMARY

. . . . . . . . . . . . . . . . . . . . . . . . . . . ii-

1. INTRODUCTION ,.......................... 11 1.1. Background ........................ 12 1.2. Conclusion ........................ 14

2. OVERVIEW OF B&W OWNER'S GROUP PROGRAM . . . . . . . . . . . . . . . 21  :

2.1. Development of New Design Basis Conditions ........ 21 2.2. Stress Analysis . . . . . . . . . . . . . . . . . . . . . . 25

3. GENERIC APPLICATION . . . . . . . . . . . . . . . . . . . . . . . . 31 3.1. Comparison of Configurations ............... 31 3.2. Plant Operations ..................... 3-3 3.3. Conclusion ........................ 34

4. REVISED DESIGN TRANSIENTS FOR THE SURGE LINE ........... 41 4.1. Instrumentation of Oconee Unit 1 ............. 42 4.2. Thermal Stratification .................. 43 4.2.1. Plant Temperature Measurements .......... 44 4.2.2. Processing tile Measurements . . . . . . . . . . . . 45 4.2.3. Development of the Prediction Correlations .-. . . 4-6 4.2.4. Application of the Correlations and Conclusions . . 48 4.3. Thermal Striping ..................... 4-9 4.3.1. Selection of the Battelle Stri 4-9 4.3.2. Data Checks . . . . . . . . . ping ...........

Data . . .... 4-11 4.3.3. Interface Calculations .............. 4 12 4.3.4. Correlation of Striping Calculations ....... 4 13 4,3.5. Conclusion .................... 4-14 4.4. Review of Operational Histories . . . . . . . . . . . . . . 4-15 4.5. Development of Revised Design Basis Transients ...... 4 16 4.5.1. Heatup Transients . . . . . . . . . . . . . . . . . 4-18 4.5.2. Cooldown Transients . . . . . . . . . . . . . . . . 4 23 4.5.3. Other Transients ................. 4 25

5. P! PING ANALYSIS . . . . . . . . . . . . . . . . . . . . . . . . . 51 5.1. Structural Loading Analysis . . . . . . . . . . . . . . . . 51 5.1.1. Choice cf Computer Program ............ 5-1 5.1.2. Building of Mathematical Model .......... 51 5.1.3. Non Linear Temperature Profile .......... 52 5.1.4. Verification Run for Displacements ........ 54 5.1.5. Structural Loading Analysis for the Thermal Stratification Conditions .s ............. 5-4 5.2. Generic Stress Indices for the Surge Line Elbows ..... 5-6 5.2.1. Purely E astic Finite Element Analysis ...... 56 C

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5.2.2. Elasto Plastic Finite Element Analysis . . . - . . . 57 4 5.2.3. Calculation of a Generic Stress Index C2 for Secondary Stress in the Surge Line Elbow . . . . . . 57 5.3. Verification of NB 3600 Equations (Equations 12 and 13, and Thermal Stress Ratcheting) . . . . . . . . . . . . . . . . 5-8 ,

i 5.4. Development of Peak Stresses ............... 5 10 5.4.1. Peak Stresses due to Fluid Flow . . . . . . . . . . 5-10 Peak Stresses due to Thermal Striping . . . . . . .

5.4.2. 5 11 .'

5.4.3. Peak Stresses Due to the Non Linearity of.the

Temperature Profile . . . . . . . . . . . . . . . . .. 5 12 S.S. Fatigue Analysis of the Surge Line. . . . . . . . . . . . . 5-13 5.6. Fatigue Analysis Results for the Surge Line . . . . . . . . 5 16

6. N0ZZLE ANALYSES . . . . . . . . . . . .-. . . . . . . . . . . . . . 61 6.1. Pressurizer Surge Nozzle . ................ 6-1 6.1.1. Geometry . . . . . ............... . 62 6.1.2. Description of Loadings . . . . . . . . . . . . . . .

62 6.1.3. Discussion of Analysis . . . . . . . . . . . . . . . 62 6.1.4. List of Assumptions / Inputs Used in Analysis . . . . 64 6.1.5. Thermal Analysis . . . . . ... . ._. . . . . . . . 65 6.1.6. Stress Analysis . . . . . . . . . . . . . . . . . . 67 6.1.7. Summary of Results and Conclusion . . . . . . . . . 69 6.2. Hot Leg Surge Nozzle . . . . . . . . . . . . . .-. . ... . 6 10 6.2.1. Geometry . . . . . . ................ 6 10 6.2.2. Description of Loadings . . . . . . . . . . . . . . 6-10 6.2.3. Discussion of Analysis . . . . . . . . . . . . . . . 6 11 6.2.4. List of Assumptions / Inputs Used in Analysis . . . . 6 14 6.2.5. Thermal Analysis of Axisymmetric Loads 6 15 tO 6.2.6. Stress Analysis of Axisymmetric Loads . . . . . . .

6 17 6.2.8. ASME Code Calculations . . . . . . . . . . . . . . 6 20 6.2.9. Summary of Results and Conclusion . . . . ._. . . . ,

6 20

7.

SUMMARY

OF RESULTS ........................ 7-1

8. BASES FOR THE B&WOG ANALYSIS - FOR PLMT SPECIFIC APPLICATION . -. . 81

9. REFERENCES ............................

9 1-

10. DOCUMENT SIGNATURES . . . . . . . . . . . . . . . . . . . . . .-. 10 1 APPENDICES A. Surge Line Data Acquisition at Oconee Unit 1 . . . . . .. . - . . . . . . A1 B. Utilization of Oconee Data for B&W Owners Group Program . . . . . . B l' C. . Supplementary Striping Information ................ C-1

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List of Tablet Table p,g, 3 1. Surge Line Insulation Comparison ................ 35 4 1. Surge Line Design Basis Transient List ............. 4 29 4 2. Design Basis Heatup Transient Definitions . . . . . . . . . . . . 4-31 4 3. Events Affecting Surge Line flow for Plant Heatup and Cooldown . 4 32

51. Specific Total Patigue Usage factors for the Surge Line . . . . . 5 17 6 1. Base Cases, Stress .ine 10. Inside Surface, Safe End to Elbow Weld .............................. 6 22 6 2. Linearized Stresses. Transient lAl ............... 6-24

63. Maximum Stresses, T ransient l Al . . . . . . . . . . . . . . . . . 6-26 6 4. Thermal Stress Base Cases for the Nozzle to Surge Line Weld . . . 6 28 6 5. Stresses and fatigue Usage for a Typical PV at the Nozzle to Surge Line Weld . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6 29 A 1. Oconee Stratification Test Signal List ............. A5 C 1. HDR Test Series TEMR PWR: Ranges of Conditions and Conditions of Test 33.19 ........................... C-17 C 2. Test Conditions . . . . . . . . . . . . . . . . . . . . . . . . . C 18 C 3. Phases of Test 33.25 . . .................... C 19 List of Fioures Figure Page

21. B&W Owners Group Surge Line Thermal Stratification Program Key Elements ............................ 2-7 2-2. Maximum Pressurizer to RC Temperature Differences imposed by Appendix G Limits . . . . . . . . . . . . . . . . . . . . . . . . 28 3 1. Lowered loop Surge Line Configuration . . . . . . . . , . . . . . 3-6

32. DB 1 Surge Line Configuration . . . . . . . . . . . . . . . . . . 3-7 4 1. Plant Data taken at Oconee Unit 1 during HPl Check Valve Test . 4 34 4 2. Thermocouple Locations at Oconee Unit 1 . . . . . . . . . . . . . .

4 35 4-3. October 89 Oconee Heatup, Wall Temperatures at Location 11 ... 4 36 4-4. October 89 Oconee Heatup, Top to Bottom Temperature Differences . 4 37 4 5. Striping Conditions . . . . . . . . . . . . . . . . . . . . . . . 4 38 4 6. Test Loop . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4 39 4 7. -Thermocouple Placement at heasurement Cross Section CC ..... a-40 4 8. Inside Wall Temperature Between 90 and 95 Degrees, Cross Section CC ............................... 4 41 4 9. Inside Wall Temperature Fluctuations 9 100 Degrees, Cross Section CC ............................... 4-42 4 10. Heatup Design Basis Transient Temperatures .......... 4-43 4-11. Development of Heatup Design Basis Transients . . . . . . . . . . 4-44 4-12. Plant Heatup Transient lAl, Temperatures ............ 4 45 4 13. Determining the Number of Events for Each Heatup Transient ... 4-46 4 14. Cooldown Design Basis Transient Temperatures .......... 4-47 p

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51. Surge Line Mathematical Model . . . . . . . . . . . . . . . . . . 5 18

52. Comparison of Displacements, X direction ............ 5 19 5 3. Comparison of Displacements, Y direction ............ 5-20 5 4. Comparison of Displacements, Z direction ............ 5 21 5 5. Finite Element Model of Surge Line Elbow ............ 5 22 6 1. Geometry of Pressurizer Surge Nozzle .............. 6 30 6 2. Finite Element Model of Pressurizer Surge Nozzle ........ 6 31

63. Example of Typical Transient .................. 6 32 6 4. Summary of Thermal Boundary Conditions (Pressurizer Surge Nozzle) 6 33 6 5. ...... 6 34 6 6. Location Structure 1ofBoundary Delta T Conditions Values (Pressurizer (PressurizerSurge Surge Nozzle)

Nozz le) . . . . 6 35

67. Location of Stress Classification Lines (PNssurizer Surge Nozzle) . . . . . . . . , , . . . . . . . . . . . . . . . . . . . 6 36 6 8. Geometry of Hot Leg Surge Nozzle ................ 6 37 6 9. Finite Element Model of Hot Leg Surge Nozzle .......... 6 38 610. Example of a Typical Transient for the Hot Leg Surge Nozzle . . . 6 39 6 11. Location of Delta-T Values (Hot Leg Surge Nozzle) . . . . . . . . 6 40 612. Hot leg Surge Nozzle Temperature Contours (F) for a Typical PV . 6 41 6 13. Location of Stress Classification Lines (Hot Leg Surge Nozzle) . 6-42 8-1. Recommended Surge Line Operational Limit ............ 83 A 1. Oconee Data Acquisition Hardware Configuration ......... AB A-2. Thermocouple Locations at Oconee Unit 1 . . . . . . . . . . . . . A9 A 3. Displacement Transducer Locations at Oconee Unit 1 ....... A 10 B 1. Oconee Unit I data taken 2/10/B9 ................ B7 B 2. HPI Flowpath - BWST to RCS ................... B8 B 3. Oconee 1 Surge Line Data: Heatup. 02/89 ............ B9 C 1. Frequency of Occurrence Versus Amplitude ............ C 20 O C-2. Frequency of Occurrenco Versus Amplituda for Hear Wall Temperature Fluctuations .......................... C 21 C 3. Striping Conditions . . . . . . . . . . . . . . . . . . . . . . . C 22 C-4. Test loop . . . . . . . . . . . . . . . . . . . . . . . . . . . . C 23 C-5. Thermocouple Placement at Measurement Cross Section CC ..... C 24 V -vi-

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O i - 1. INTRODUCTION i

This report summarizes the -B&W Owners Group program addressing the technical issues described in NRC Bulletin 8811 (reference 1). The analyses described in '

this report confirm that all surge line pressure boundary components (including

all nozzles) satisfy applicable code stress allowables for the operating BtW lowered loop plants

Arkansas Nuclear One Unit 1, Crystal River Unit 3, Oconee Unit 1, Oconee Unit 2, Oconee Unit 3, and Thrae Mile Island Unit 1, considering

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the effects of thermal stratification and thermal striping. The Davis Besse Unit 1 (raised loop plant) surge line is currently being evaluated and will be addressed in a future supplement to this report.

In light of the extensive technical program developed to. respond generically to NRC Bulletin - 8811, .the. Babcock and Wilcox Owners- Group formed a Thermal O stratification Workin. oroon. The Worki# Groon incioded representatives from each nember utility. The represenatives were chosen for their expertise 'in thermal-hydraulics or structural mechanics. - The : Working- Group has been extensively-involved in this program and has functioned as technical ' overseers, technical coordinators between the utilities and BWNS. and program managers.

The introduction provides the background for the thermal stratification, striping l and cycling issues and a summary of the surge line fatigue- analysis results and L conclusions. The remaining sections of the report are as follows:

o l . Section 2 describes the technical approach which has been developed by the B&W Owners Group,

. Section 3 discusses the justification for the generic approach taken by the lowered loop plant owners,

. - Section' 4 describes the development of - the new- design basis thermal-hydraulic conditions for the surge line,

.. Sections 5 and 6 describe the stress and fatigue analyses performed for 1-i 1

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i the surge line piping and its nozzles,

. Section 7 provides conclusions resulting from the B&W Owners Group program with regard to new design basis transients which represent surge line thermal conditions and the structural integrity of the surge line,

. Section B states the conditions which form the basis for the analysis.

. Section 9 lists all references, and

. The appendices provide a detailed discussion of the striping and stratification tests which were used as aids in developing the design basis transients.

1.1. Backaround The surge line in B&W 177FA plants contains approximately 50 feet of piping which connects the pressurizer lower head and the reactor coolant hot leg piping.

During plant operation, the reactor coolant system (RCS) is pressurized with a steam bubble in the pressurizer. Thus, the pressurizer contains saturated fluid while the remainder of the RCS is subcooled with temperatures generally at least 50F cooler than the pressurizer fluid. The surge line provides the means by O which the pressurizer accommodates changes in RCS inventory. Thus, RCS fluid flows through the surge line during surges into and 'out of the pressurizer.

During RC pump operations, there is an outflow from the pressurizer due to l pressurizer bypass spray flow.

Due to differences in density, the reactor coolant can stratify in the horizontal piping section whereby the fluid temperature varies from top to bottom with the warmer fluid located above the denser (cooler) fluid. This phenomenon, known as thermal stratification, is most pronounced during outsurges from the pressurizer.

During an insurge or outsurge under stratified conditions, a phenomenon known as thermal striping may occur at the fluid layer interface. . Thermal striping is' a rapid oscillation of the thermal boundary interface caused by interfacial waves and turbulent effects. The original surge line fatigue analyses performed for the B&W 177FA plants assumed that uniform temperatures existed at each surge line location and did not account for thermal stratification which cause: additional bending moments in the piping. The original surge line fatigue analyses also did not account for thermal striping which affects the fatigue usage at the inner 1-2 4

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l t L considered in the original stress ane.iysis. The effects of this bowing have been-

- observed in the surge line of the Portland General Electric Company Trojan plant during each refueling outage since 1982. The piping deflections have ~ been ,,

l observed to have resulted in reduced clearances at the pipe restraints and, in some instances, contact of the piping and the pipe restraints has' been observed.

Similar occurrences have also been noted at Beaver Valley Unit 2.

In order to confirs pressurizer surge line integrity, the Nuclear ~ Regulatory Comission issued NRC Bulletin Number 8811, Pressurizer Surae Line Thermal ,

Stratification (December 20,1980). This bulletin requires certain actions of licensees of all operating pressurized water. reactors J(PWRs). The actions  ;

applicable to the case at hand are paraphrased below.

la. At the' first available cold shutdown after receipt of the bulletin,.

and which exceeds seven days, conduct a visual inspection of the pressurizer surge line.

Ib. Within four months-of recaipt of the bulletin, licensees of plants in operation over ten years are requested to demonstrate that the .

pressurizer surge. line' meets ~ the applicable design codes and other FSAR and' regulatory commitments for the licensed life of the plant,-

considering thermal stratification and. thermal striping in the fatigue and stress evaluations.. .(For licensees of plants which have been in operation less than ten years, this action must be completed j

within one.' year of receipt of the-bulletin.) Or provide'the staff with a justification for. continued operation while- a detailed l

analysis of the surge line is performed by implementing items le and Id below.

Ic. If- necessary.- obtain plant specific surge -.line thermal and displacement data. Data can be obtained through' collective efforts if sufficient similarities in' geometry and ' operation can be demonstrated.- ,

L Id. Update the fatigue and stress analyses to ensure compliance with the ,

i applicable Code requirements.

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A portion of the B&W Owners Group program was presented to the Nuclear Regulatory Commission Staff on September 29, 1988 and April 7.1989. An interim evaluation, BAW-2085 dated Pay 1989, provided the staff with a justification for near term operation for all of the operating B&W 177FA plants (reference 2). The NRC concluded that sufficient information had been provided to justify near term operation for B&W plants until the final report could be completed (reference 3).

This report summarizes the entire B&W Owners Group program and documents compliance with action items Ib, Ic, and Id of NRC Bulletin 8811 to demonstrate code compliance for the 40 year life of the surge line.

1.2. Conclusion The surge lines for the six B&W lowered loop plants are shown to fulfill the 40 year design life. The structural analysis of the surge line and associated nozzles has accounted for thermal conditions (thermal stratification, thermal striping, and thermal cycling) existing during the life of the-plants. The highest cumulative usage factor for 40 years of operation occurs in the vertical elbow located at the bottom of the riser (elbow B on Figure-31). This elbow has been calculated to have the largest cumulative usage' factor for the Oconae Unit 2 O >#r9 14 ith < '# or o > < r the ao 1 r e 49 iir o6 o heatup/cooldown cycles. For the non' Oconee' lowered loop plants, the highest

>$o cumulative usage fact'o r occurs in the same elbow for Arkansas Nuclear One Unit' 1 (ANO 1) and is equal to 0.66 for the 40 year design life and based on 240 heatup/cooldown cycles.

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2. OVERVIEW OF B&W OWNER'S GROUP PROGRAM Revised surge line design basis transients have been developed based on an assessment of information from all the B&W plants including detailed thermal stratification data taken at Oconee Unit 1, thermal striping data collected by Battelle Karlsruhe, other plant heatup, cooldown, and operational data, and plant operating and surveillance procedures. The number of occurrences for each design basis transient has been determined for each plant from operating histories and a fatigue stress evaluation has been performed for the surge line of each plant.

The B&W Owners Group Materials Committee has developed a comprehensive program to address all technical concerns identified in NRC Bulletin 8811. The program is divided into two basic sections. The first addresses the need to understand relevant thermal hydraulic phenomena occurring in the surge lines, to develop O operational recommendations, as appropriate, to reduce the severity and impact of thermal stratification in the surge line, and lastly to develop a new thermal-hydraulic design basis for the structural analysis and fatigue evaluation of the

! surge line and the associated nozzles. The design basis thermal transients reptuent the final product of the first part of the program. The second part l

of the program addresses all structural analyses required to assess the integrity of the surge line and associated nozzles for the balance of the design life of each of the plants.

This report is a summary of the entire program. Figure 1 shows the relationship of the key elements of the program, which is further described in the following two subsections.

2.1. Development of-New Desion Basis Conditions The thermal-hydraulic phenomena which had to be accounted for in the surge line were thermal stratification, thermal striping, and thermal cycling. These phenomena occur to- some degree in almost. all modes of plant operation. As a A

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result, the surge line conditions had to be carefully considered from cold shutdown through heatup, power escalation and cooldown.

Operational events et power also cause surge line thermal cycling although these cycles have been shown to be relatively inconsequential as a result of evaluations performed in this program. This is primarily due to the small  !

difference in temperature between the pressurizer fluid and the RCS hot leg fluid.

Thermal cycling is a transient phenomenon and is thus associated with mass or temperature changes in the RCS. Thermal stratification can occur in the surge >

line only during moderate to low flow rates through the surge line and iray exist l in a steady state as well as in a transient condition. Thermal striping requires the existence of thermal stratification. Quantitative treatment of these  ;

phenomena requires knowledge of the following:

1. Previous operating experience (surge line operating experience),

2. Operating practices that influence surge line conditions,

3. Thermal stratification, and

4. Relationship between thermal striping and thermal stratification.

Previous operating experience includes the total of all thermal cycles and

' stratification conditions that ht.ve existed since plant startup. Data does not exist in order to describe the complete history of the surge line transients in the plants. Therefore, two things have been done to estimate-the operational  ;

history. First, each utility reviewed plant' operating logs to identify each

-heatup and cooldown performed since plant startup. The utilities then retrieved I data sets for the overall RCS parameters for the heatup and cooldown events to typify thesi evolutions. These data sets have provided the bases for key j operational characteristics such as the average times for the various phases of l- a' plant heatup, the frequency and magnitudes of surge line flow transients, and ,

the relationship between the pressurizer and reactor coolant loop temperatures.

To understand surge line thermal hydraulic service conditions, the Oconee Unit I surge line was thoroughly instrumented to record the thermal transients in the l

surge line for plant heatup, power escalation, full power. operation, and plant I cooldown. Surge line displacement instrumentation was also added to all major-lO 22 1

f sections of the Oconee Unit I surge line. The Oconee data collection effort, described in more detail in Appendix A, has provided circumferential temperature measurements at several different axial locations along the surge line in ,

addition to displacement measurements for each major displacement axis. These data have been used to develop a correlation of surge line temperature responses to the surge line flow rate for given end point pressurizer and RC loop temperatures.

In parallel with the Oconee Unit 1 msasurement program and the collection of historical plant data, a review of the operating procedures for all plants has _ ,

been performed to better understand those plant evolutions likely to cause surge line upsets. This review has been supplemented with interviews of plant operators at each site. Evolutions causing significant surge flows have been identified and quantified, within the limits of the plant data. Plant data indicated many more surge line upsets than could be described and quantified as determined from the procedures. Hence. the mechanistic surge line events have been augmented with random, or unexplained, surge line flow events. The random events have been based on typical plant data for changes in pressurizer level during heatup and cooldown operations.- The total: number of surge li'ne ' flow occurrences have been based on plant data. The results of the overall effort j have provided the bases for assumed surge line upsets that take place during

! heatup, cooldown,-and power operation.

An important part of the operating procedure review has dealt with t"e potential upper bound for the pressurizer to hot leg temperature differei :- The - B&W Owners chose to relate the' potential upper limit on the. surge ina end point

. differential temperatures to pressure limits- imposed according to 10CFR50, Appendix G (see Figure 2 2.). The Appendix G limits have been reviewed for each plant and a bounding composite . pressure / temperature (P/T) limit has been developed to cover all plants for the 0 to 2 Effective Full Power Years (EFPY).

period, the 2 EFPY to 1990. time period, and the period from 1991 to the end of life. Since no B&W designed lowered loop plant transient has been identified to have approached or exceeded the pressure / temperature limits at or near cold RCS ^

conditions (50F-120F), these limits represent an upper bound on the pressyrizer temperature to hot leg temperature difference. The survey of plant heatup and cooldown data has shown that- none of the recorded transients approached the 2-3 4

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k pressure temperature relationship adopted for this program. The maximum theoretical surge line top-to bottom temperature difference is 397F, as discussed in Section 4.5.1.2. The maximum surge line top-to-bottom temperature difference measured at Oconee Unit I was 280F.

The third item on the list of key elements addresses thermal striping. A review of the literature on thermal striping has been made to identify potentially useful experimental data for application to the surge _line analyses. As noted in the interim submittal (reference 2),_ most experimental work reported in the literature applus to fluid conditions very different from those that exist in a PWR surge line. The majority of work- has been done at low pressures and .

relatively low temperatures. Some of the experiments used different concentrations of solutions to simulate the density gradients _that can exist at higher fluid temperatures. One set of experiments, those performed by Battelle-Karlsruhe Laboratories, stood out as clearly superior in simulating the PWR surge line conditions. BWNS performed a detailed review of published reports on the experimental apparatus and partial test results and has followed this review up with direct contact with the principal investigator, Dr. Lothar-Wolf. The B&W Owners Group decided to purchase the raw data from Battelle for the series of g)

(_. nine test runs made at typical PWR surge line conditions. Details of the experimental setup and tested conditions are addressed in Section- 4.3. The important rescits that have been derived from the Battelle dsta are the relationship of thermal striping amplitude and frequency to the pipe fluid cenditions. The thermal striping data correlation allows the determination of striping characteristics for any given surge line flow rate and imposed top-to-bottom temperature difference.

The product of the thermal-hydraulic prooram is a revised set of surge line design basis transient descriptions that account for thermal' cycling, thermal stratification, and thermal striping. All design transients considered in the previous' design basis for the surge line have been modified to account for all three thermal phenomena. All design basis transients involving surges have been considered in the evaluation.

Results of the thermal-hydraulics part of the program consist of the input f

• the stress analysis of the surge line itself, the associated nozzles at each enu, and the one-inch diamete drain nozzle connection at the bottom of the lower OV 2-4

s. . - _ _ _ _ - _ _ _ _ _ _ _ _ _ _

a

, horizontal run. The stress analysis portion of the program is described in the

\V next subsection.

2.2. Stress Anal n.11 The first phase of the stress analysis involved building a structural mathematical model containing the pressurizer, the surge line, the hot leg, the reactor vessel and the steam generator. This structural mathematical model has been verified by using the measured surge line temperature data from the Oconee Unit I heat up of February 1989 to predict surge line displacements. These predicted surge line displacements agree well with the measured surge line displacements (see* subsection 5.1.4 and Figures 5-2 through 5-4).

The structural loading analysis has been performed using the new thermal-hydraulic design basis. This generates the internal forces and moments which have been used for the stret , analysis of the surge line and the nozzles associated with its endpoints.

The applicable piping code is the 1986 Edition of ASME Code NB-3600, in accordance with NRC Bulletin 88-11 which states: " Fatigue analysis should be p performed in accordance with the latest ASME Section III requirements V incorporating high cycle fatigue". A Code reconciliation has been performed for each plant to remove all concerns associated with.using the later Code (1986 Edition) for the fatigue stress analysis of the surge line. For the Code reconciliation, each plant's surge line stress report has been reviewed (the code of record for piping analysis is not the same for all the surge lines).

Using finite element analysis for the elbows and simplified equations elsewhere in the surge line, all stress intensity values (Equations 12 and 13, and Thermal Stress Ratchet) have been found to be within their allowables (taken from ASME Section III, Appendix I). Therefore, the elaste-plastic fatigue analysis has been performed in accordance with NB-3653.6(c). To account for the thermal-hydraulic conditions defined in the new design basis, the surge line fatigue analysis includes thermal stratification, pressure ranges between the thermal stratification conditions, thermal striping, fluid flow and temperature changes leading to through-wall temperature gradients, and the additional imlized stress due to the non-linearity of the top-to-bottom temperature profile.

()

V 2-5 l

l

in the NB 3600 elasto plastic fatigue analysis, all applicable surge line O. locations have been analyzed, including the drain line nozzle considered as a branch connection. The design basis transients are genere11y the same for each one of the six B&W lowered-loop plants. Tho number of occurrences of those transients, however, is unique to the plant to be analyzed due to differences in operating history. The elasto plastic fatigue analysis has been performed and applied to each of the six plants being analyzed. For each one of the six B&W lowered-loop plants, the total cumulative usage factor is less than 1.0 at all surge line locations.

In addition to the structural analysis of the surge line described above, detailed stress analyses of the pressurizer and hot leg nozzles have been performed to demonstrate compliance with the ASME Code, Section 111.

Finite element models have been made of both nozzles and the thermal and pressure stresses have been calculated using the revised design basis transient descriptions as input. Piping loads acting on the nozzles have been taken from the structural analysis of the surge line and have been combined with the pressure and thermal stresses. Stress and fatigue analyses were performed in accordance with the requirements of the 1986 Edition of the ASME Code, Section Ill, NB-3200 and NB-3600. The analyses demonstrate that the cumulative usage factor for each nozzle is less than 1.0.

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2) Portions of the Appendix G hmits are less restrictive; than the original O 2EFPY limits due to improvements in the analytical methods used to develop the 5j limits. In spite of this, the Appendix G hmits are expected to become more restrictive in the future. Consequently, the maximum pressurtzer to RC temperature differences are expected to decrease.

4 2-8

r O

3. GENERIC APPLICATION Factors affecting surge line thermal stratification have been assessed .to determine if the B&WOG pl 1ts can be evaluated _ generically. The assessment.

addressed two different types of factors: (1) those.that are inherent in the equipment design and 72) the plant specific operating and surveillance procedures that may influence the surge line condit 'ons. The following two subsections summarize these evaluations, 3.1. Comoarison of Confiaurations A review of the surge line piping for all B&W 177 fuel assembly plants shcws that, with the exception of Davis-Besse Unit 1 (081), all plants have the same nominal dimensions and configuration. The lowered loop and raised _ loop (DB-1) r n surge lines are shown in Figures 3-1 and 3-2 respectively.

The surge line for each configuration is constructed of 10" diameter schedule 140 austenitic stainless steel pipe (inside diameter 8.75") with a wall thickness of 1". The surge lines are insulated with a reflective / mirror insulation having similar characteristics for all B&W plants as tabulated in Table 3-1.. The end nozzles connect the surge line to the hot leg and to the pressurizer. In the lower horizontal piping run, a one inch = diameter- nozzle made of austenitic material connects a drain line to the surge line. The drain nozzle itself (dimensions and material) is the same for all 'six lowered loop surge lines. The plant specific location of this drain nozzle is indicated in _ Figure 5-1 of this report.

With the exception of THI-1, snubber restraints are positioned to restrain the surge line in case of a seismic event, and allow for pipe movement during design thermal cycles. The snubbers are placed at different locations along the surge line at each plant. As long as the snubber displacement is within the range of free travel, the effect on surge line stresses is negligible. The TMI-l surge line does not contain any seismic snubbers, restraints, or supports.

3-1

. . . -. - -- . - .. - ~ ~ - -

At Davis Besse Unit 1, a fixed pipe whipf restraint construction is used atLeight A At each whip restraint'_ location, an impact g locations along the surge line.

collar acting as - a spacer -is- affixed to the _ pressurizer surge line. Free movement of the surge line is determined by preset gaps between shims, applied to the inside of the _ whip restraint, and the pressurizer surge line collar. ,

Eight pipe whip restraints, one spring hanger, and three snubbers provide support -

for the surge line under th6 various_ design loading conditions.-

Surge line pipe hangers are used only:at the Davis Besse and Crystal River plants (these hangers are variable spring hangers).- The surge line piping for the other B&W 177 plants is free hanging. Crystal River has three- dead weight hangers' located on-the surge line. The available travel in the hangers is sufficient to prevent each one of these hangers from bottoming out._ Therefore, the hangers have a-negligible impact on the surge line stresses.

Other than nozzles at the pressurizer and the' hot -leg connections, the only other-piping connection on the surge line is a one-inch drain. nozzle located =in the-lower horizontal section. During the heatup, cooldown, and power operation of a lowered loop plant, ~ this drain pipe is valved off and has no. effect on-the ,

j ,

surge line response to surges.

Based upon an evaluation of the surge line1 configurations _and plant operations,

- the lowered loop plants. are sufficiently-similar to be evaluated generically.

However the differences between Davis-Besse Unit'l and the lowered loop plants l'ed to the decision to install special instrumentation at Davis-Besse to gather data during the heatup from the 6th refueling _ outagelin:the summer of 1990.

Considerations leading to this decision are the following:

. The Davis-Besse surge line configuration differs.significantly from-the

! lowered loop plants. The lower horizontal. run is somewhat shorter, and-there is an upper. horizontal run in-excess of 20 feet compared to.the 2.5 -

foot length in the-lowered loop plants.

. The " qe line at Davis Besse incorporates eight fixed pipe whip restraint structw es, .with impact collars clamped to the surge line. - These' impact-collars interrupt the _ insulation,: permitting ' gaps to'n either side. The:

increased heat loss -potentially affects the degree of stratification.

3-2 t' e 9 -

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. At Davis-Besse' the power operated relief valve (PORV) inlet condensate drain is connected to the surge line drain upstream of the drain isolation valve. Condensate reflux into the surge line depends upon' heat lusses from the line, and could have some influence on the surge line stratification response.

The data from the special instrumentation at Davis-Besse are being used as the l basis for a plant specific stress and fa;1gue analysis which will be submitted as a supplement to this report.

3.2. Plant Operations The plant operational aspects which' affect the magnitude and number of thermal 1 l cycles applied to the pressurizer surge line have been evaluated in order to l formulate the design basis cycles.

During the evaluation, the applicable plant operating procedures and plant data <

have been reviewed. In addition, plant operators have been interviewed at each pl ant. The results indicate that all of the B&W plants basically operate in a l similar fashion with some minor differences. The focus of the operational review .

l

- has been the _heatup/cooldown and initial RCS pressurization phases, since the potential for significant thermal stratification conditions exists only during these phases.

During power operating conditions and during operating conditions wherein-RCS temperature is near " Hot Standby" (approximately 530F) all of the plants operate in a similar fashion and the thermal stratification potential (pressurizer saturation temperature minus RCS hot-leg temperature) is relatively small.

During design basis: event transients, the transients imposed on the surge line are virtually identical for all of the lowered loop plants.

Since the reactor vessel operational P/T limits, in accordance with 10CFR50 Appendix G, provide the upper limit of the pressurizer surge-line thermal stratification potential during the heatup/cooldown and initial pressurization i

phases, and these limits are a function of the effective full power years (EFPY) of operation, the magnitude of the thermal stratification gradients as well as l' the actual number of heatup/cooldown cycles have been grouped on the basis of the O

Q 3-3 e ~v-- +-

periods of the applicable Appendix G limits. Actual plant data have shown that operation of the B&W plants has been below the reactor vessel P/T. limits.

Plant data from the instrumented Oconee Unit I surge line as well as the

• installed pressurizer surge line thermocouple and pressurizer level data from each of the other B&W plants have been utilized to define the number and magnitude of thermal stratification cycles for the generic design basis.

Based on the P/T path taken by each of the plants during past heatups and cooldowns, the magnitude of future thermal stratification cycles was developed and has formed the basis for evaluating future surge line fatigue.

The design basis thermal stratification cycles developed for this evaluation are generically conservative in t,oth number and magnitude. Recognition of the operational actions which produce such cycles will reduce the actual future cycles, in number and magnitude, to less than those which have been evaluated.

3.3. Conclusion The lowered loop plant configuration and plant operations are quite similar and a generic development of design basis transients is justified. The number of occurrences of each type of generic transient will vary based on each plant's operating history.

Davis-Besse Unit 1, which is a raised loop B&W plant, requires a plant specific analysis due to the differences discussed in Section 3.1. The analysis for Davis-Besse Unit I will be addressed in a supplement to this report. The methodology described in this report is generally _ applicable to the Davis-Besse Unit I analysis. This includes the correlation of stratification and striping, the synthesis of design transients, the structural modeling techniques, the structural loading analysis, and the fatigue analyses of the surge line and its associated nozzles. Differences from the material contained in this report, due to plant specific structural and operating conditions will be identified and justified in the supplement containing the analysis for Davis-Besse Unit 1.

3-4

. .- - 7 Table 3-1. Surge Line Insulation Comparison PLANT < INSULATION TYPE: THICKNESS HEAT LOSS

' MFG, ' BTU /(HR-FT^2)

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4. REVISED DESIGN TRANSIENTS FOR THE. SURGE LINE The purpose of NSSS design basis transients is to provide conservative thermal-hydraulic conditions for representative plant transients- that are used in the structural analysis and fatigue evaluation of the reactor coolant system components.. Transients described in the design basis include normal - plant evolutions (such as heatup and cooldown), upset conditions (such as reactor trips),emergencyevents(suchas_rapidRCSdepressurization_followingasteam:

generator tube . rupture), :and faulted events (such as a steam line break).

Depending on the classification of the transient, various- stress criteria must be met and demonstrated in the stress report for the NSSS components. For normal-and upset transients, i.e. those that will occur during the normal course of plant operation, the fatigue of each NSSS component must be evaluat'ed. The-design basis transients provide- the thermal-hydraulic input for the fatigue evaluations.

Typical parameters that are specified for each design basis transient include the applicable temperatures and pressures during the transientL condition. Other parameters, such as the flow rate, are also specified to enable the determination of heat transfer coefficients from the fluid to the metal wall. In the case of thermal stratification, several additional parameters are of. interest to' the stress analyst. These include:

. Temperature difference from top.to bottom of the surge .line,

. Changes in the magnitude of stratification,

. Top-to-bottom cross-sectional temperature profile,- and

. Location of the temperatureL interface between stratified fluid layers.

4-1

.i

-The developnent of the new design basis transients involved the following:

(O> b Measurement of surge line data at Ocunee Unit 1,

2. Development of surge line stratification correlations,

3. Development of thermal striping correlations,

4. Review of operational histories, and -

5. Formulation of new design basis transients.

Each of these areas is addressed in subsequent subsections.

4.1. Instrumentation of Oconee Unit 1 The original Reactor Coolant System instrumentation package in each B&W-designed -

plant included temperature detectors in the hot and cold legs of the- reactor coolant loops, in the lower half of the pressurizer (i.e. the liquid space), and.

on the outside of the surge l'ine. The single detector on the surge line is a thermocouple mounted on the lower horizontal run. The surge line thermocouple, generally mounted on the side: of the pipe,- is for use in monitoring for gross changes in surge line temperature such as those that might exist if the

^ pressurizer spray bypass flow is stopped. However,.this thermocouple measurement -

is not adequate for use in determining the existence of stratified conditions in the surge line. ,

Following comparisons of the as-built dimensions of the surge lines it was concluded that a single plant could be instrumented to-provide typical-data for thermal response under conditions of thermal stratification. Oconee Unit I was selected as the plant to be instrumented.

The objectives of the instrumentation program have been-to determine:

. The magnitude of the thermal stratification including the maximum top-to-bottom piping temperature differential,

. Variations in the thermal stratification with axial position along the surge line,

. The changes in surge line displacement that rt._ lt from thermal stratification,

. The plant operctions that' cause thermal stratification cycles, and 4-2

_-_________.__________m_._.____-._._m__ _ _ _ _ . .

. The temperature response of the surge line to pressurizer level O

y changes.

To meet the above objectives, a comprehensive instrumentation package was prepared including 54 thermocouples mounted on the outside circumference of the surge line and 14 displacement instruments affixed to various parts of the line.

The thermocouples and displacement instruments were connected to a data acquisition system allowing continuous monitoring of all instruments. The system also recorded numerous permanent plant computer signals which have provided data for the determination of changes in RCS conditions that affect the surge line flow rates. Details of the instrumentation and the data acquisition system are included in Appendix A.

The instrumentation package and data acquisition system were installed during the

~

refueling outage of January 1989. Data were recorded as-the plant prepared for and went into its normal heatup in early February. There was no interference with normal plant operations and no changes to procedures were made to accommodate the data acquisition or to reduce the effects of potential thermal stratification. Data were recorded throughout the heatup, power escalation, and for several days at full power.

The Oconee data clearly show that thermal stratification develops after the pressurizer is heated in preparation for starting the first reactor coolant pump.

Thermal stratification cycles of varying magnitude occurred throughout the heatup and power escalation. Figure 41 is a sample of the data from the February 1989 heatup. The correlation between pressurizer level changes and the surge line temperature response is easily seen. The quantification of this and other relationships is the next step in the preparation of revised design basis transients. This quantification has been viewed as an essential element of the overall program since no two plant evolutions will occur in exactly the same way.

A means to predict the surge line conditions for any given event, based on the RCS parameters and plant conditions, is needed. The next subsection addresses the development of the surge line correlations based on Oconee Unit 1 data.

4.2. Thermal Stratification Correlation Stratification correlations have been developed from plant surge line temperature measurements recorded at Oconee Unit 1. The plant measurements cover a wide 4-3

range of system conditions and evolutions, including power operation, heatup, cooldown, pump starts and stops, and spray. Stratification correlations have been developed for each of the key surge line locations. These correlations have been based on observed temperature responses to changes in surge line conditions; they: have been adjusted as necessary to . ensure that the predictions are conservative as well as realistic. The temperature differences predicted:by these correlations have been compared to the observed temperature differences.

The same coding used to generate these predictions has also been used to analyze the synthesized plant transients, to produce revised ~ design basis thermal transients for the surge- line analysis. The following sections _ provide additional information- regarding_ the plant measurements, the development and refinement of the ' prediction-correlations, and their performance.

A number of assumptions are implicit in the development of the stratification correlations. Principal among them is the application ~ of' outside-wall temperature measurements to' predict inside-wall temperatures. The time constant of the surge line pipe wall is on the order of four minutes (and varies with the

.onvective heat transfer coefficient). Therefore, rapid temperature fluctuations occurring at the inside surface are attenuated at the outside measuring surface.

O (Indeed, the high-frequency fluctuations associated with thermal striping are not observable at the outside wall.) The extreme temperature changes of greatest.

interest in stress analyses generally develop over relatively long . periods, equivalent to many pipe wall time constants. Therefore, the application of' outside measurements to predict inside-wall stratification is acceptable.

4.2.1. Plant Temoerature Measurements l Plant surge line temperatures have been recorded at nine cross sections spaced throughout the line, as shown in Figure 4-2. At each of the six_ instrumented cross sections-in the horizontal piping runs, from 5 to 10 thermocouples were

]

distributed circumferential1y to provide complete coverage of temperature profile across the height of the pipe. Temperature measurements were recorded at either 20-second or 1-minute intervals.

4-4'

The following three sets of data have been selected for this analysis:

1) Power operation of 19 Sept 89,

2) Cooldown beginning 22 Sept 89, and

3) Heatup beginning 02 Oct 89.

These data sets span.a wide variety of plant conditions and evolutions. The power-operation data include a reactor coolant pump trip. The cooldown data span a period of three days of operations, it includes pump trips and the actuation of both regular and auxiliary pressurizer spray. The third data set, the October heatup, spans approximately one week. The plant was initially cold, with no pumps running. Data recording continued through the intermittent operation of the reactor coolant pumps, heatup using pumps, and power escalation. A period of pressurizer recirculation is also included with pressurizer heaters manually energized and pressurizer spray control in automatic.

The acquired plant data has been extensively cross-plotted and compared. The temperature distributions at each cross section, the sequential response of temperature versus location in the surge line, and temperature responses to insurges and outsurges, confirm the sensitivity and self-consistency of the surge O' line temperature measurements.

4.2.2. Processina the Measurements The measured t' aratures at instrumented cross sections have been used to-l estimate the elevation of the current, local, average temperature, as well as the elevations at which the local temperature reaches one-fourth and three-fourths of the differential from the lowest local temperature to the highest. These elevations provide indications of the location and vertical extent of the thermal interface. The top-to-bottom temperature difference:has also been obtained as a function of time.

The top-to-bottom temperature differences at each measurement section in the horizontal piping runs and time have been compared to determine the overall maximum temperature difference versus tima (the vertical piping runs were observed to be approximately isothermal). With the exception of the short horizontal run' at the hot leg, the . local temperature differences among the several measurement locations have been within 20% of each other. The 4-5 7  ? ?" w

(*-w-=y9 y

temperature difference at location 11, approximately midway in the horizontal run nearest to the pressurizer, has generally been observed to be slightly larger-than the temperature difference at _ other locations. Therefore, the top and bottom temperatures at location 11 have been used for subsequent comparisons of predictions to observations.

The overall maximum observed top to-bottom temperature difference has been processed further to identify the most limiting stratification cycles. First, the times and temperatures at which the time-rate-of change of-- the maximum-  !

temperature difference changed sign have been identified- These arrays"of extreme temperature differences and times have been analyzed using the Ordered Overall Range method of cycle counting developed by Fuchs et al (reference 4).

For a specified threshold: temperature _ difference, the Ordered Overall _ Range method has identified those large and small- temperature differences (commonly referred to as " peaks" and " valleys", or "PVs") which are significant in terms of stress. These results have been compared to the significant PVs which have-been predicted rather than measured, as discussed in the following subsection.

4.2.3. Develooment of the Prediction Correlations Correlations have been developed to predict plant surge line temperatures versus time using the following global plant conditions:

. System pressure (or saturation temperature)

. Hot leg temperature

. Pressurizer level (or surge line flow rate)

. Reactor coolant pump operating status

. Spray valve status A prediction correlation has been developed for each temperature of interest.-

As mentioned previously, the top-of-pipe and bottom-of-pipe . temperatures at location 11 (see Figure 4-6) have been of primary interest for stratification in the horizontal piping. Predictions have also been developed for the measurement-locations nearest the hot leg-to-surge line nozzle and the pressurizer nozzle, to provide estimates of the fluid temperatures at these nozzles.

4-6

{

The measured surge line temperatures often respond to.an insurge or. outsurge in O a decaying exponential fashion. They have also exhibited threshold surge line  ;

V flow rates below which they are unresponsive, .and time delays before they begin '

to respond. -Therefore, the prediction correlations have consisted of the corresponding four temperature characteristics, as follows:

. -Threshold,

. Time delay,

. Final- temperature, and

. Time constant.

The threshold flow rate and time- delay are; self-explanatory. The final-temperature and time constant are combined to describe the decaying exponential temperature change' from the initial temperature to -the: final temperature;- the - i time constant is the elapsed time required for the fractional _ temperature change d

. to approach the total temperature change by a factor of (1-e ).

The prediction correlations also identified the " mode" of current surge line condi.tions -- insurge, outsurge, or bypass. Spray and pressurizer-recirculation modes are also assigned. The bypass mode-is the default mode, and is used when O- no other mode of surge line conditions is currently active.

The prediction correlations have been developed by examining brief' periods of -

data when the observed surge line temperatures appear to be responding to a single or predominant event, such as the outsurge associated with a reduction of j pressurizer level. The apparent time constant and finalitemperature' have been-

- estimated.- These parameters have been estimated for several intervals, at a variety of plant and surge line conditi_ons, and have been combined to form the

' basic prediction correlations. These correlations are based on calculateo surge

]

line flow rates obtained from changes in pressurizer level and do not require knowledge of specific operations occurring intthe plant.

Preliminary correlations based on. identifiable discrete events have been refined -

and adjusted.-in two steps. First, the predicted temperatures have been compared to their observations. The temperatures predicted were adjusted-as required to-ensure the predictions are realistic but conservative. The prediction-of the  ;

top-of-pipe. temperatures have been skewed to render them slightly more responsive t 4-7

}-

9 -- ,w t e~

l than observed, and the bottom of-pipe predictions have been adjusted to make them slightly cooler, and less responsive, than observed.

(Vl The final adjustments have been- based on the predicted versus observed significant temperature differences.' As was done for the observed (maximum) temperature differences, the top-to-bottom differences of the predicted temperatures have been processed to yield the extreme differences (PVs), and the differences which are significant in terms of stress (the significant PVs). The predicted and observed significant PVs have been compared over each of the data sets and evolutions. The prediction correlations have been adjusted, if necessary, to yield conservative but realistic predictions of the significant PVs. The hot / cold fluid interface elevations have been predicted and compared to the observations in a similar fashion; the predicted interface elevations have been skewed towards the centerline of the pipe'for conservatism.

Figures 4-3 and 4-4 demonstrate the predicted and observed temperatures and temperature differences at location 11, during the period containing the overall extreme PV. ,

4.2.4. Anolication of the Correlations and Conclusions t O The stratification correlations have been used to process the synthesized plant l ()

transients, described later in Section 4.5. The coding used to compare the predictions to observations has also been-used for processing the synthesized plant transients to generate revised design transients for surge line analysis.

I The stratification correlations provide predictions which are appropriate for surge line stress analyses. The stratification correlations have been developed from plant data spanning a wide variety of conditians and evolutions. The correlations are responsive to basic surge line conditions, rather than to the

'The 'significant" temperature differences in the comparisons of predictions to observations were the results of the Ordered Overall Range method of cycle counting. This method identified those PV-pairs which were of the proper time sequence and which defined a temperature change greater than the specified threshold. For the comparisons of predictions to measurements, this threshold was set to obtain a number of PVs suitable for comparison. When the predictions wera applied to synthesized plant transients, as described later herein, the counting threshold was set to identify those PVs which were of significance to stress or fatigue. The fatigue analysis nas shown that thermal cycles occurring with a temperature difference greater than 50F must be accounted for.

g 4-8 Q

- specific type of plant evolution. . Finally, the' stratification correlations  ;

provide realistic but conservative estimates of the extreme surge line i

stratification conditions .

4.3. Thermal Strioina -

Surge line striping has been estimated using a-correlation of experimentally-observed striping and is based on - the "TEMR-PWR" experiments performed at Battelle Institute, Karlsruhe, FRG. These experiments have been performed at plant typical fluid conditions and used large diameter, horizontal, insulated, metal pipe which is characteristic of the horizontal piping runs of PWR surge  :

lines. The acquired experimental data has been self-consistent; mornver, it provided adequate temperature versus time definition of' the high frequency, localized thermal-hydraulic interactior,s which characterize striping. The temperature data have been processed to determine interface characteristics as well as striping frequencies and amplitudes. The interface information has facilitated the interpretation of the observed striping variations among tests.

The maximum striping amplitude has been correlated with.the governing fluid conditions, as has been the distribution _of striping frequency versus amplitude.

These striping efforts are described below; Appendix C'provides a summary of the O striping literature and a description of the Battelle experiments.

M 1. Selection of the Battelle Stricina Data .

There are four major sources of striping information:

. Boiling Water Reactor (BWR) feedwater nozzle _tes'a ,

. Liquid Metal Fast Breeder Reactor (LMFBR) tests,

.- Argonne National- Laboratory (ANL). tests, andl Battelle-Karlsruhe tests (HDR).-

The BWR feedwater nozzle striping tests results; (reference 5) were quite extensive. The - associated studies demonstrate the ability to combine low-temperature test data with high temperature test data and with plant striping-data. The application of these results illustrate' the use of probability distribution functions.- But' the geometry and hence the striping interaction of the feedwater nozzle are not like those- of. the surge line, therefore these results are not applicable to the surge line.

h 4-9

___--__--_c_ . . . - - . - - -

l Woodward (reference 6) has investigated: striping and stratificat' ion in a 1/5-scale model of an 1.MFBR. The temperature fluctuations in transparent, horizontal ]

pipes have been examined using dye and.thermocouples. Results _were based on -

inside -diameters of 4 and 6.5 inches, .and- hot __ and cold fluid temperatures of approximately 130F and 70F. Woodward references- the experimental- work _ofL j Fujimoto et al (reference 7) for a fluid to wall heat transfer coefficient. The Fujimoto tests also pertain to low temperature ' conditions, but . employ the addition of calcium chloride to the warmer fluid stream to obtain plant-typical j

differences -of fluid densities. (The fluid temperatures.were used simply as-tracers of the streams of differing densities.)--

Kasza et al, at ANL (references 8_through-10), have conducted extensive tests of.-

stratification and striping.- Combinations of horizontal and. vertical piping runs ~

have- been tested, including bends in both the vertical: and horizontal: planes, i These tests also employed the relatively low bounding fluid temperatures'of the' ,

tests previously mentioned. Kasza-et al_ have published a limited. amount of power-spectral-density information for striping; the reported signal energy peaks between0.1and0.6Hz,anddecaysexponentiallywithincreasing[ frequency.

Wolf et al (references 11 through 15) have also conducted extensive experimental l

l studies of stratification and striping in the;HDR facility at Battelle Institute, Karlsruhe, FRG. Unlike the other experimental - offorts, the HDR tests were performed in a large diameter, insulated, metal pipeLusing plant-typical fluid The pipe was extensively instrumented with. fast-response conditions.

thermocouples.

l The thermal-hydraulic conditions governing striping (buoyancy', inertia, and viscosity) are the appropriate bases on which to compare the several striping tests discussed above. High-frequency fluctuations of wall temperature, namely, striping, are due to - rapid undulations /of. the fluid thermal . interface -- the relatively narrow zone between fluids of dissimilar temperatures and densities.

These undulations are commonly caused by the interactions among. the fluid

- )

buoyant, viscous, and inertial forces. Buoyancy tends to -stratify the fluid, whereas the fluid inertia and viscosity tend to mix the fluid and'to disrupt stratification. _These competing effects give rise to interface instabilities. ,

4-10

l l

l The interacting fluid forces are conveniently quantified in terms of two l

O eime s4enless raties of fleie <erces. 18ese are the aex eles ##m8er. 8ich is the ratio of inertial to viscous forces, and the Grashof number, which is the ratio of buoyant to viscous forces. A third dimensionless ratio is commonly used to gauge the presence and significance of stratification, and has also been extensively referenced in studies of interface instabilities. This is the Richardson number, which is the ratio of buoyant to inertial forces.

The LMFBR, ANL, and HDR tests previously mentioned have been compared on the bases of these governing force ratios, in Figure 4-5. The Reynolds number is readily varied experimentally, but the Grashof number is more. intractable. It includes the difference between the bounding (hot and cold) fluid temperttures, and the fluid thermal expansion coefficient. In a low-pressure facility, the maximum temperature difference is limited. Similarly, the thermal expansion coefficient of water is a strong function of temperature; it increases almost threefold going from 200F to 500F. The low pressure facilities are again limited in this regard. Thus the Grashof numbers of the low-pressure tests are far.below those obtained at plant-typical conditions. Only the HDR tests by Wolf et al lie.

in the appropriate range. For this reason,-the HDR data has been selected for ,

use in this investigation of surge line striping.

4.3.2. Data Checks Thermal stratification and striping in a horizontal line were examined by Wolf et al in the "TEMR" series of tests. Multiple temperature measurements have been made- in a 20 foot long horizontal section of 15.6-in inside diameter pipe (Figures 4-6 and 4-7). Three subseries of TEMR tests were distinguished by differing obstructions placed at the junction of the pipe with the reservoir of '

hot fluid. The junction was left unobstructed in the "PWR" subseries of tests;

these tests are considered to be the most relevant to the surge line geometry, therefore the "PWR" tests have been examined.

The complete test measurements have been obtained. These consisted of hundreds of measurements processed at 10 Hz, for tests ranging up to 1000 seconds in l duration.- The acquired data has been subjected'to rudimentary tests as it has been transferred to the analysis computer. Each temperature reading has been scanned for out-of-range measurements, and each signal has been checked to verify 4-11

that it did not remain constant ' throughout a test. Offending temperature measurements have been set to negative infinity, to prevent their processing and p* lotting.

The vast majority of the instruments and readings have passed these checks. The few temperature readings and instruments which failed these tests have generally corresponded to those which had been identified as suspect by the experimenters,_ '

in their test report (reference 16). Two sets of temperature measurements contained out-of-range readings, both in . Test 33.14. The earlier failure occurred just after 640 seconds of an 872.7-second test, therefore data processing has' been interrupted at 640 seconds with little net. impact on the ability to draw conclusions:from this test.

Each of the tests was initialized with the pipe hot .and isothermal, providing a convenient cross-ch'e ck of all the temperature measurements. The tests were initiated by introducing cold water at the bottom of one.end of the horizontal pipe -- the responses to this cold stream provided another straightforwai4 check of thermocoup1- response. Finally, the thermocouples were spaced sufficieni.ly closely in the circumferential direction that multiple adjacent thermocouples as p often responded to the passage of an individual- interface -undulation, V demonstrated in Figure 4-8.

The temperature meast.rements consisted primarily of thermocouple triplets --

these measured the temperatures of the fluid ne.ar the wall, of the inside wall, and of the outside wall. 'Several cross sections of the pipe were instrumented.

At the most-heavily instrumented cross section, the thermocouple triplets were placed every 2-1/2 angular degrees near the mid-height of the pipe, from 90 to 105 degrees, increasing to 30-degree spacing at the top and bottom of the pipe.

4.3.3. Interface Calculations This investigation of striping depended on- the response of individual thermocouples, and hence on there being a temperature measurement within the thermal interface, which is the region of steep temperature gradients and maximum striping potential. The striping calculations have been preceded by an examination of the characteristics of the thermal interface, in particular'the-proximity of this interface to the installed thermocouples.

O 4-12

. - . - ~y ,.

The thermal interface has been characterized using three fractional temperatures, O T l " , 'T2 " , and ' T3 " . Fractional temperature T1 has been defined as the coldest local fluid temperature plus one-fourth of the difference between the hottest and coldest local temperatures; similarly, T2 has been defined to be one-half of the way from the coldest to the hottest temperature, and T3 has been defined to be three quarters of the way. The definition of these fractional temperatures is somewhat arbitrary -- they could have been assigned other fractional values, for example. They have simply been defined as convenient intermediate temperatures.

The elevation of the average temperature (T2) is approximately located at the center of the interface,- and the difference between the elevations of T3 and T1 l indicated the vertical extent-of the (half-) interface. The elevations of the

( fractional temperatures have been extracted from time-smoothed data.

l At two measurement cross-sections and each time-of measurement, each group of measurements from three adjacent thermocouples which bracketed a fractional temperature has been curve fit using a second-order relation of temperature to (thermocouple) elevation. This relation has been solved for the elevation of the l corresponding fractional temperature. The elevations "Z1", "Z2", and "Z3" of the corresponding fractional temperatures demonstrated interface development and stabilization in each test. Comparisons of the interface location and thickness among the tests have permitted their correlation in terms of the governing fluid l forces. More importantly, the elevation of the stabilized interface has .been compared to the thermocouple elevations, to identify those tests in which the observed striping most closely corresponded to the actual maximum striping.

l 4.3.4. .lorrelation of Strioino Calculations-The recorded temperatures provided good definition of the temperature trends, but-slightly underestimated the extreme temperatures. The actual extreme l temperatures are estimated by- solving for the zero-slope point of the second-order fit to the three temperature measurements bracketing the apparent extreme.

l l The results of these estimates are illustrated in Figure 4-9.

The Ordered Overall Range method of Fuchs et al (reference 4) has been used to count striping cycles. Counting has been performed separately for each inside-wall thermocouple at the three instrumented cross sections. Counting focused on-the relatively ~ st.able period which followed the initial realignment of O 4-13

y i!

i temperatures and-interface development;: The stablefperiod has-been;. counted both; i

. overall and ' bM f halves, :tof provide: a gauge;of self-consistency. For each -

~

instrument and counting period, the countingLthreshold has been decremented from; a the highest threshold, which just passed the' difference between the most extreme.

' _ temperatures throughout? the period, -down to a~ minimum ' threshold. - L The. cycles'  ;

1 counted at each threshold, divided by the counting interval; formed the so-called? '

" cumulative frequency of occurrence".LThis is the rapidityfat I which' temperature; changes larger than:the threshold have been observed; - The distribution of these  ;

cumulative! frequencies of occurrence versus1 striping amplitude formed readily--

-correlated relations.. '

LThe maximum striping = amplitude <for each test hasfaean compared and correlated -

with the governing fluid conditions.1The maximum stripingfamplitude.of each test l has beenLindexed - according to the ' proximity of theltherma.1? interface;to.1 the:

measurement' thermocouples. ; Generally, the;1argest :stripingi amp 1.itudes.Loverj ai j' range of governing conditions have been observed whenithe interface bracketed a~

thermocouple, and'vice versa. The correlation of,maximumistriping;: amplitude d t versus governing- conditions has been constructed .toLreproduce th_ese Llargest-

-amplit'ude points. The inaximum. striping' amplitudes oflthe final correlation have -

. been increased by 100 (of4the imposed temperature? difference)J to allowifor-residual uncertainties. - Finally,Vthelfrequenci versus-amplitude distributions'-

L obtained byi Kasza et -al (reference -10) < have been c consultedL to check-- the t correlated trends at the highest frequencies; the decreasing amplitudes at. higher L frequencies-of-occurrence have'_been confirmed by, the-_ distributions:obtained by -

Kasza et al.

-4.3.5. Conclusion .

Stratification and striping information have been extracted from the-HDR data, .

.and correlated with the governing fluid forces. The HDR data is singular in that-

- it 41one. was o taken at.: conditions. typical of theiplant surge 111ne. :The '

rxperimental data has been found to be self consistent.1The HDR data was taken overn a: fairly broad range -of fluid _ conditions. Moreover, 'the:: striping-correlations'have been based on the ratios of governing fluid 1 forces, rather than fluid -properties. Therefore, these correlations provide 1usefullestimates of

$ striping over the expected range of-plant evolutions and surge line conditions- .

4-14

...+.a.

. . - . . . ..ge..,. e -- s-,s.,--- . , -,-a-.w. g % , ,,_wg,-m, , , , , g, W,.,,, ,,,4w s u + ye r t *m 9-- v 4 w w 'P- evtP+ v'w w +- + v -- w'e m -w *gy V

The striping correlations can be used with confidence to predict striping effects

, ,) in the plant surge lines.

(

4.4. Review of Operational Histories >

The revised design basis transients are intended to represent the operational history of the plants and to bound future transients. Hence, past operetional information is required to generate revised transients. An information base which includes plant operating data, operating procedures, surveillance procedures and operational limits has been collected from utility and BWS records. In addition, discussions with plant operators have provided records of first hand experience.

Representative operational data for sixteen heatups and fifteen cooldowns have been supplied by the plant utilities. The parameters provided inc N 2:

RC Temperature (Hot leg and Cold leg)

RC Pressure Pressurizer Temperature i

Pressurizer Level Makeup flow rate Letdown flow rate Surge Line Temperatures (original thermocouple)

I, s')

RC flow RC spray valve status (open/close)

Steam Generator Pressures Steam Generator Levels The data for these parameters provide the general bases for describing heatup and cooldown design basis transients.

In conjunction with the plant data review, operation and surveillance procedures have been reviewed to identify events that might cause stratification cycles.

In deciding whether or not generic design basis transients are appropriate, operational procedures have been reviewed for differences. It has been determined from the surveillance procedure review that valve testing involving RCS inventory changes has been the most significant type of surveillance procedure in terms of thermal cycling of the surge line. As a result of this review, Transient 22 of the design basis transients has been revised to account for HPI check valve testing. The operating procedures have been used to determine the major flow events to be accounted for in the heatup and cooldown

,o

() 4-15 l

1 j

transients. Random flow events have also been added to ensure that all pressurizer level changes observed in the plant data are taken into account.

All applicable pressure-temperature (P/T) limits and requirements have been collected to determine the bound of operation for heatups and cooldowns. These pressure-temperature limits and requirements are comprised of 10CFR50 Appendix 0 limits, decay heat removal system (DHRS) limits, Net Positive Suction Head (NPSH) requirements for RC pumps, minimum sub-cooling pressure requirements, and fuel compression limits. The actual pressure temperature history indicated by the historical heatup/cocidown data and the Appendix G limits has been used to create both representative and bounding pressure-temperature relationships for

! the heatup anu cooldown transients, l 4.5. Develooment of Revised Desian Basis Transients The revised surge line design basis include the transients listed in Table 4-1.

It is the redefinition.of the surge line conditions for these transients that forms the bases for reanalysis of the structural integrity of the surge line piping and its associated nozzles.

The greatest impact of stratification and striping occurs during conditions when the temperature difference between the pressurizer and the RCS hot leg is large.

The largest temperature differences occur during plant -heatup and cooldown.

These transients contribute most of the cumulative fatigue usage for the surge line piping. Therefore an extensive effort has been made to develop detailed descriptions for these plant evolutions. The design basis plant heatup and cooldown transients have been completely redefined in this program. Other transients included in the design basis have generally been retained in terms of the existing surge line boundary conditions' (pressurizer temperature, RCS j temperature and flow rates), but thermal stratification and striping have been included in the surge line transient descriptions.

l

[ Typical variation of RC hot leg temperature as a function of time has been determined for heatup and = cooldown transients based on plant data. The l-pressurizer pressure variation is based on pressure-temperature (P/T) relationships, both the 2 EFPY P/T limit and available plant data. Pressurizer temperature has been developed as a function of RC temperatures from the P/T relationships.

4-16' U  :

1 l

The Oconee design basis differ from the current design b.'is for other plants in O ** t the ocomee de ie" e 4, 4 c'#de dirr re t of this evaluation, these heatup rate differences have been disregarded and the t a r te - rer th a#rne -

time for heatup has been set at one conservatively long period for the heatup events. The various heatup rates given in the design basis are important for the major components, i.e., reactor vessel and steam generator, but are not l

necessarily conservative for description of the surge line conditions.

Operational events that affect surge line flow during heatup and cooldown have l been identified by a review of plant data, operating procedures, and surveillance test procedures for the lowered loop plants. Surge line flow rates resulting from these operational events have been characterized using available plant data, l

l analysis, and engineering judgement. All operational events which affect RCS inventory, volume and spray flow are included. The number of surge events per i plant transient and time between occurrences have been ' estimated for each significant type of event affecting the surge line flows. Random surge line flow j events have been included to ensure that the total number of surge events corresponds to the number based on plant data.

Composite heatup and cooldown transients with an appropriate sequence of events, ,

timing, and urge line flows have been generated. The total number of surge line flow events for each type of transient has been described and categorized according to severity in terms of pressurizer to RC temperature differences.

Probable past operations and the projected improvements in operations affecting surge line conditions have been taken into account in defining the heatup and cooldown transients. The correlations discussed in Section 4.2.2 and 4.2.3 have been used to generate the surge line thermal response to the flow events.

The number of occurr:nces for each type cf heatup and cooldown transient are based on operational records and on the total number of heatups and cooldowns stated in th, original design basis. The original design basis assumptions for the number of occurrences for the other transients has generally been retained.

The following sub-sections describe the detailed development of the- heatup, coldown, and other events.

4-17 l

. - . . . _ . - - . - . . _ . . . , . _ . . _ _ _ _ _ , _ _ _ _ _ . , , _ , _ _ . . _ - - , _ , . , . - _ , ~ , _ , _ . . . - . ,

b  !

. +

i 1

'4,5.1. Heatuo Transient 1 ,

I 4.5.1.1. Heatuo Transient Descriotions and Wagher of Occurr s Heatups are categorized into five transients (Transients IAl - IA5) in order to- '.

conservatively represent plant heatup transients occurring 'during_ the 40 year life of the plant. The RC and pressurizer temperatures for these transients are r

] shown in Figure 410. A heatup transient is defined as the operations from a RC j

I temperature ~of 70F-to_8% power. Thefirstthreetransients-(IA1,1A2,'and1A3):

represent past operations and the last-two transients-(IA4 and:lA5)-represent-

future operations. The surge line flow events occurring during.the heatup have been defined to be identical _ for all' five types of transientst_ however, _the RCS l &nd pressurizer temperatures vary. .Thei development 'of these; transients is

represented by the schematic in Figure 4rli and is discussed below. l Transient IA1. This transient-has been- conservatively defined.to bound past _ ,

plant operatiens in terms of the maximum possible pressure in z

the' RCS. The resulting surge line temperatures are shown.in FigureL4312. For these1 heatups, core decay . heat has been-assumed to heat the RCS to r 120F priori to the operator 4

O chi vi . ** Rr r c rx < r >t rti . ac R ma-Pressurizer temperature is based on the pressure temperature j;

relationship defined by the 2 EFPY Appendix G 1_imit~.

Transient 1A2 ~The RC temperature.for this= transient'r aresents Leatups in

which decay heat-has not been availabit / or initial hectup and the RC pumps were started with-an RC temperature of 70F. The pressure temperature relationship for this -transient-i . represents the fraction of plant heatups-which have occurred f with' a higher than- average RC ~ pressure 1 forr given' RC temperatures based on plant historical data.

Transient 1A3 This. transient represents typical plant heatups for the past '

l operating history of BWlowered loop plants. - Decay heat has been' assumed' available 'for? these heatups. Land the RC t

temperature has ' been assumed to. be' 120F at the time ~ the I.

pressure required to start.the' first RC pump has been reached.

l .

O :4-w l

. l i

e The pressure temperature relationship for this transient is representative of historical plant data.

Transient 1A4 The RCS temperature as a function of time represents the heatups for which decay heTt will not be available for the initial portion of the heatup. The pressure temperature relationship has been based upon plant data in which the pressure required for the first RC pump start was reached at an approximate RC temperatste of 70F. The pressure-temperature limits discussed in Section 8 will help ensure that this basis remains valid for future heatups.

Transient lAS The RCS temperature as a function of time represents the heatups for which decay heat will be available for the initial portion of the heatup. The pressure temperature relationship has been based upon plant data in which the pressure required

! for the first RC pump start was reached at an approydmate RC temperature of 120F. The pressure temperature limits i specified in Section 8 will hel.p ensure that this basis remains valid for future heatups.

Some plants do not start RC pumps until a higher RC temperature than those represented in the transients (>120F) is obtained. The pressurizer to RC l-temperature difference for these heatups is usually lower due to a lower system pressure for RC temperatures. To conservatively represent all lowered loop B&W plants, the transients are described with RC. pump starts at RC temperatures at 70F and 120F.

1 The number of occurrences for each type of heatup transient, tabulated for each lowered loop plant in Table 4 2, has been determined using utility records and l plant data as the schematic shows in Figure 4 13. The number of Transient 1Al I events has been obtained by taking a conservative estimated fraction of the total .

number of heatup events occurring over the first two years of operation and an equivalent number of heatups that occur. during hot functional tests. The remaining fraction of events occurring nece the first two years are grouped with the events that have been distributed between Transients 1A2 and 1A3 based on plant experience to date. Based on plant data, approximately 15% of these-past 4-19 e

9

-_.-,_.m r_.,%.. , _ _ . _ , _ _ _ , , , , _ . , ., , , _ . _ , ,, _ . , _ _ . . , . , v..

l l

heatup events are categorized as type 1A2. The remainder of events are included g)

( in type 1A3. For future events, approximat'aly 15% are categorized as type 1A4 and the remainder of design events are included in type 1AS.

4.5.1.2. Maximum Pressurizer-to RCS Temp 3rature Difference A maximum stratification temperature differential has been specified for a small fraction of the total number of design transients to account for operating conditions in which the system pressure corresponds to the maximum allowed by the 2 EFPY P/T curve (Transient 1A1), it has been assumed that the RCS temperature reached approximately 80F at the tirae the pressurizer temperature reaches 480F (saturation temperature for the pressure limit of 565 psia at RC temperatures below 275F). The resulting tempert4ture differential between the pressurizer and the RC loop of 400F is considered to be a realistic, conservative upper bound for this temperature difference. Although lower RC loop temperatures are possible during typical heatup operations, it is considered very unlikely that the simultaneous occurrence of lower RC loop temperatures with RCS pressures approaching the RCS P/T limits has occurred at any plant. The maximum value of stratification (surge line top to bottom temperature difference) that resulted p

v for the revised surge line design transients is 397F.

4.5.1.3. Boundary Temoeratures as a Functh n of Time she RC temperature versus time for the revised heatup transients is similar to the existing transients contained in the original-design basis except that the duration of the heatup has been extended to be more representative of actual plant operations. Temperature holds have been added and heatup rates adjusted to give reasonable agreepent between the heatup for analysis and the available plant data. The pressurtzer temperature versus time plots for Transients 1Al through 1A3 have been developed based on the relationship of_ pressurizer temperature to RCS temperature corresponding to available plant data. Transients 1A4 and 1A5 represent future operations based upon B&W Owners Group l

recommendations for plant operation discussed in Section 8.

l The time for heatup is based on octual plant experience, which shows average heatup times of approximately 60 hours6.944444e-4 days <br />0.0167 hours <br />9.920635e-5 weeks <br />2.283e-5 months <br />. The surge line analysis conservatively

[D 4-20

assumes each heatup requires 75 hours8.680556e-4 days <br />0.0208 hours <br />1.240079e-4 weeks <br />2.85375e-5 months <br />. The time durations for each portion of the heatup have been developed from available plant data for the following:

RC loops cold (<100F), pressurizer hot (>200F)

Heatup RCS heldoperations (RCS at constant temperature temperature increase) throughout a ll modes Hot, zero power Power Escalation 0 to 8%

4.5.1.4. Surce line Flow Rates for Heatuo Changes of flow of pressurizer or hot leg fluid through the surge line leads to thermal stratification transients. Unlike the RCS and pressurizer temperature relationship, direct operator guidance does not exist for pressurizer level control Procedures and operations have changed over the years with some changes decreasing the number and magnitude of surges while others have increased this number, it is not possible to describe every plant event that influenced the surge line flow rate and consequently the effects on the thermal transients for the surge line piping and nozzles. However, by supplementing the major flow events with random flow events, a typical heatup transient ,as h been generated which rem,istically represents the number and magnitude of surge flow events.

Each heatup transient case (Transient 1Al - 1AS), has been assumed to have the same flow events. On the average, these transients' represent the typical l operations occurring durirg the' life of a lowered loop plant. The five heatup transients differ in the relationship between RCS and pressurizer temperature as discussed in Section 4.5.1.1.

l Typical plant heatup operations that may affect the makeup flow to the RCS are listed in Table 4 3. The operations that have been judged as.significant and quantifiable are considered the major events which have been accounted for in the transients. RCS temperature changes affecting RCS inventory are accounted for in addition to events such as RC pump starts, certain surveillance tests, and venting operations. The response of makeup flow rate to each of these events is accounted for in the development of the transients.

Transient surge line flow events found in Oconee and other plant data from unidentifiable causes have been incorporated in the design transients as random events. The available plant data has been statistically analyzed to characterize (3 4 21 V  ;

the random flow events in terms of flow rates and changes in pressurizer level.

h To accomplish this, the following steps have been taken:

1) Plant heatup data have been evaluated to determine the average and standard deviations for pressuri7er level changes (inches) and flow rates during operation over various ranges of RCS operating temperatures. Also, the average number of flow events has been detennined for the heatup operations.

2) The average number of flow events has been adjusted so that the number of events per unit time is consistent between the prescribed heatup length and the data.

3) The number of random flow events has been determined by taking the total number of flow upsets derived from the plant data less the number of defined flow events.

4) Flow change data have been treated as a normally distributed, random variable and divided into three ranges based on the normal distribution curve. Flow rates representing the three regions have been sequentially repeated for the level events in the synthesized transients, b 5) Random flow events have been unifonnly spaced over the appropriate time corresponding to RCS operating temperature ranges.

Pressurizer spray operation for the B&W plants is classified in one of three modes: pressurizer RCS boron equilibration, automatic actuation of spray and bypass spray. The bypass spray operation is identical for all B&W plants with a four pump operation bypass flow rate between 1.5 and 5 gpm. Surge line flow rates are accounted for based on the number of pumps operating.

The operation of full spray differs between the Oconee Units and the other three B&W lowered loop plants. The original modulating spray valves have been replaced with on off acting Target Rock spray valves for the Oconee Units. All other plants have operated their entire life with the original modulating spray valves.

When the spray is actuated for heater / spray operation to bring the pressurizer boron into equilibrium with the RCS boron concentration, the Target Rock solenoid operated valve cycling results in periodic spray flow, whereas the modulating valve allows the spray flow to be held constant. The Oconee Target

(

l l

(3 / 4 22 l

Rock valves supply approximately 315 gpm maximum flow rate and the typical modulating valve supplies 190 gpm maximum flow rate.

Spray operation occurs for the following two operations as part of the design basis heatup transients: (1) pressurizer-RCS boron equilibration occurs at hot s; stdown conditions (Tavg - 532F) and (2) spray cycling occurs during the increase from 0 to 8% power due to the increase in Tavg causing the spray setpoint to be reached.

Although the Oconee spray valve and the modulating spray valve types do differ, their differences do not have an impact on thermal stratification. When the spray is actuated automatically, the difference in flow rate between the Oconee Target Rock spray valve and the standard modulating type of valve is not significant in terms of stratification since both flows are large enough to flush the surge line of any stratified fluid. The action of the spray f_ low in flushing the line of stratified fluid also effectively eliminates the conditions that can lead to the occurrence of striping. Automatic actuation of the spray valve in either case. Target Rock or modulating valve, rid the line the conditions required for striping to exist.

4.5.2. Couldown Transients l 4.5.2.1. Cooldown Transient Descriotions and Number of Occurrences Cooldowns are categorized into two transients in order to conservatively represent RC cooldowns from 8% FP to refueling. temperatures for the 40 year life l of the plant. The RC and pressurizer temperatures for these two transients are shown in Figure 4-14. The time duration and the surge flow events are the same for both transients, but different nressure temoerature relationships are used.

All lowered loop plant cooldown transients are represented by the two transients discussed in the following paragraphs.

I V 4-23 t

m w

r l

Transient 101 The pressure temperature relationship for this cooldown

' transient has been based oh the least restrictive Appendix G limits and available plant data. The least restrictive Appendix G Ifmits (highest allowed RC pressure for a given RC temperature) are used to determine the pressurizer temperature '

4 for RC temperatures below 200F and above 400f. It has been determined by a review of plant data that the plants operated well below the cooldown Appendix G limits for RC temperatures ,

between 200F and 400F. The pressurizer temperature has been assumed to be 250F above the RC- temperature for this RC temperature region as this bounds the available plant data.

Transient 182 The pressure temperature relationship for this transient it.

identical to Transient 181 except for the intermediate (200F<Trcs<400 ) pressure temperature relationship for which representative data is used instead of the bounding plant data.

Based on plant data, approximately 15% of the past cooldown events have been ,

I /~N ca teaorized as type 1B1. The remainder of events have been included in type 182.

U Future cooldowns have also been categorized with 15% as type 1B1 - and the remaining 85% as type 1B2.

4.5.2.2. Boundary Temocratures as a Function of TIER The average time required for cooldowns, from plant dats experience, indicates an average time for a complete cooldown of 46 hours5.324074e-4 days <br />0.0128 hours <br />7.60582e-5 weeks <br />1.7503e-5 months <br />. The surge line analysis assumes that each cooldown transient requires 75 hours8.680556e-4 days <br />0.0208 hours <br />1.240079e-4 weeks <br />2.85375e-5 months <br />. This long duration conservatively allows surge line thermal events to stabilize. The time durations for each portion of the cooldown transient have been estimated using available plant data for the following phases of the cooldown operationst Power Decrease from 15 to 0% Power Cooldownoperations(RCtemperaturedecrease)

Cooldown hold throughout all modes RCS below 200F, Pressurizer hot Pressurizer cooldown

.The RC temperatures versus time for the cooldown transients have been based on the existing transients contained in the original design basis with the exception 4 24

- , ,m , -- -

r ,,-,,-g -v- , ,a , ---+ e , w -wm ww w -

that the times have been ertended to be more representative of actual plant operations. Temperature holds have been added and cooldown rates adjusted to give reasonable agreement between the cooldown for analysis and the available plant data. The pressurizer temperature versus time plots for the cooldown transients have been developed based on the relationship of pressurizer temperature to RCS temperature based on the available plant data.

1,L2 3. Surae line flow Ratts for Cooldog Cooldown flow events have been developed in a manner similar to that method used for determining heatup flow events as discussed in Section 4.5.1.4.

The operations that have been judged as significant and quantifiable are considered the major events to be accounted for in the cooldown transients.

Actual makeup system response to each flow event is accounted for in the development of the transients. The spray flow required throughout a cooldown for depressurization is included in the surges specified for the cooldown transients.

Random flow events have been added to ensure that the total number of flow events for the transients adequately represent actual operating experience. Available o plant data has been statistically analyzed to describe the flow events required.

O The same method which determined the random heatup flow events, outlined in ,

Section 4.5.1.4, has been used to determine the number of random cooldown events.

4.5.3. Other Transients Existing spray flow rates, temperatures, and pressures from the original RCS design basis transients have remained unchanged for most transients other than plant heatups and cooldowns. The design basis transients in which thermal stratification and striping require evaluation are listed in Table 41.

Spray flow rates, temperatures, and pressures from the original d) sign basis transients have generally been used to define the boundary conditions used to evaluate thermal stratification and striping. The numbers assigned to this transient are similar to the original design basis. Surge line flow events caused by changes in the volume average temperature in the . RCS have been generated by calculating the surge rate from the existing plots of RC temperature '

versus time contained in the original design basis transients. A few  ;

(O

(_/ 4-25 I

l

modifications and additions to the list'of transients have been made as discussed in the next few sections, f

4.5.3.1. steadv state T==aarature variations - Transient-13 l

Transient 13' has been redefined to be more realistic in tems of- the flow i L conditions experienced in the plants for normal steady state operations at power. 4 The previous' description conservatively ahd unrealistically described the steady; l state transient as an oscillation in reactor power of 1 percent and a temperature oscillation of 2F on ~ an approximate period of 6 minutes.' ' Based on the results i

, of an evaluation of actual plant characteristics from observations of experienced control room operators, the variation is' more appropriatelyi described as a variation of temperature with an amplitude of 0.5F on a 15 minute period. This transient has been arbitrarily defined to occur over a two hour period and the-number of transients have been determined accordingly for the forty year life of i

- the plant.

• 4.5.3.2. Pressurizer RCS Baron Eauilibration - Transient 200 Transient 200 has been added to' describe the coffects' of spray . and' heater * '

operations to equalize the pressurizert and RCS boron' concentrations,. This c

O transient has been described bot

• fer oReration with the on/off tv,e of s,rax valve _(Oconee Target Rock valves) and the modulating . valves.- The operation:

involves - use tof ' spray. flow through'. the : pressurizer ; to D eause' the" boron concentration to approach that in the RCS. 4 modulated spray flow;of. about 50:

gpm-has been used in the transient-description. The maximum flow rate used-for L the Target Rock valve is 315 gpm.

The operation of pressurizer spray and heaters to circulate reactor coolant for purposes'of bringing the pressurizer boron concentration to within range of'the concentration in the RCS-is perfomed. roughly twice a week; :The operation will.

normally last approximately eight . hours. The. number of L these ' transients occurring during the 40 year life of the plant has been detemined accordingly.

.4.5.3.3. HPI Safety In_iection and Check Valve Test . Transient 22 Core Flood Tank- (CFT) check valve test operations have been combined with the'  ;

heatup and- the--cooldown transient descriptions.: However, the . HPI safety; injection ~ and the check valve tests are described separately. The present O a M 7 W 3-+g tagVg .fataw e- W u"a-pg .r9Pe$ t.-(c4. t + 9y>- .h -.y.e-.+=g-y gy .wsw gwa g. v s g g wwNegy e w- gsyc,7 % ,+vage+-y>g >Nen 4 hahG .ddee e-we We e n P-V e#--wq v' b9+r-9 * + - ' -'**WW*-r&P&'

l t

I description of the HPI transient given in the original design basis transients O cdescribe rr a d> t th r ** i 3 ct4 t t 'r d <> 4 t d d t "> r tiv 'r the conditions at the injection nozzles. . An additional transient has ,

been added to describe the HPI check valve tests and the resulting effects on the surge line thermal conditions.

The HP! safety injection test and the HPI check valve test have been described

- for four different operating conditions which represent appropriate points during past and future heatups and cooldowns. The test conditi.ons for the pressurizer and RC temperatures are summarized as follows:

Safety Injection' Check Valve. RCS Temperature Pressurizer Test Test' Temperature

~

22A 1 22A 2 90F 480F 22B 1 22B 2 70F 400F 22C-1 22C 2 280F 530F 22D 1 22D 2 280F $00F Transient 220 is based on future HPI safety injection and check valve tests not being performed with a pressurizer to RC temperature difference greater than 220F.

For the safety injection tests, a flow rate of 100 gpm per nozzle (total of 400 gpm) is assumed to be established for a period of about 30 ' seconds. The HPI check valve test flow rates have been set ' to - approximate the change in pressurizer level and surge line temperature response based on the measured response at Oconee:as part of the data acquisition discussed in Appendix A. The length' of time between flow increases has been set- to give a decrea:e in temperature of about 180F at the top 'of the surge line after every' flow event, which is in agreement with the measured plant data. An offsetting outsurge flow

' of 15 gpm for a two hour period counters the 100 gpm insurge flows and maintains the final level approximately equal to the level at the start of the simulated test. The plant. data collected at Oconee Unit 1 indicated at least three flow pulses based on measured level in the pressurizer. The number of flow pulses.for analysis purposes is four for each test series.-

A

.Q 4 27 0

4

.r... r .._~, . , - . - , . ..%'., .

, ,_.,.....-,.~a,

__w.., ,_..u. ... . . c -[. e s., s _.

In estimating the number of HPI safety injection tests and check valve tests for

-the 40 year plant life, plant operators have been consulted and the surveillance requirements'have been reviewed. Requirements for some plants stipulate only ,

that tests be conducted every 18 months as a maximum, whereas other plants are required to test during shutdown conditions if the test has not been performed in the last 90 days (or if maintenance or modifications have been performed that would affect the HP1 system flow characteristics). In estimating.the number of Transient 22 events, the 90 day requirement has been used to conservatively estimate the number of events. Frequent check valve tests were not' required prior to about 1980. Therefore there have not been many of these tests during the period when the 2 EFPY P/T limit curves were in effect.-  !

, O

. l i

I l

i 4-28 Q

.I

Table 41. Surge Line Design Basis Transient List O Transient Transient Description Modification from Original ID Transients (ODB - Original Design Basis) _

lAl Past Heatups Trc of 70F Completely redefined to to 8%FP, 2 EFPY Appendix G realistically represent the limits specifies P/1 most severe heatup occurring relationship, decsy heat during the life cycle of the avail able, plant.

lA2 Past Heatups Tre of 70F Completely redefined to to 8%FP, higher than usual realistically represent heatups RC pressures specify P/T occurring with higher than relationship, no decay usual pressures for all RC heat. temperatures.

lA3 Past Heatups Trc of 70F Completely redefined to to 8%FP, P/T relationship realistically represent an based on average P/T average heatup.

relationships, decay heat available, lA4 Future heatups Tre of 70F Completely redefined to to 8%FP, P/T relationship represent future heatups for based on plant data in which decay heat is not used to

/ which decay heat is not heat RCS before starting pumps, available. _

(see path ABE in Figure 8 1) 1A5 Future heatups Tre of 70F Completely redefined to to 8%FP, P/T relationship represent future heatups for based on plant data for which decay heat is used to which decay heat is heat the RCS to 120F before available, starting RC pumps. (see path CDE in Figure 8-1) 181 Cooldown 15%FP to Completely redefined to refueling temperature, P/T represent a bounding cooldown relationship based on based on least limiting maximum RC pressures. Appendix G limits (highest pressures) and plant data.

1B2 Cooldown 15%FP to Completely redefined to refueling temperature, P/T represent an average cooldown relationship based on based on least limiting average pressures. Appendix G limits and average plant data.

2A Power Change from 0% to 15% Surge line temperatures based FP on ODB boundary conditions.

O (j 4 29 i

l

Transient Transient Description Modification from Original I i ID Transients (0DB - Original DesignBasis) __

l 2B Power Change from 15% to 0% Surge 'ine temperatures based FP on ODB boundary conditions. (

3 Power loading 8% to 100% FP Surge Line temperatures based on ODB boundary conditions.

Surge Line temperatures based l 4 Power unloading 100 to B percent on ODB boundary conditions.

5 Ten percent step load power Surge Line temperatures based increase on ODB boundary conditions.

6 Ten percent step load power Surge Line temperatures based decrease on ODB boundary conditions.

7 Step Load decrease 100 to Surge Line temperatures based 8% FP on 000 boundary conditions.

8 Reactor Trip All trips now included in Transient B. Previously, certain trips were included in other transients.

9 Rapid Depressurization SL temps based on ODB boundary q conditions.

SL temps based on ODB boundary 10 Change of RC Flow Rate conditions.

13 Steady State Temperature SL temps based on ODB boundary Variations conditions.

14 Control Rod Drop SL temps based on ODB boundary conditions.

19 feed and Bleed Operations SL temps based on ODB boundary conditions.

20 Miscellaneous Transients A complete transient was added to describe pressurizer spray and heater operations used to equilibrate pressurizer & RCS boron concentrations.

22 Test Transients Complete transients were added for the HPI check valve test.

The HPI Safety injection Test was completely modified.

4 30

Table 4 2. D6 sign Basis Heatup Transient Definitions Transient Past/ P/T Number of Events

" ' " E hy ANO 1 CR 3 00 1 OC 2 0C 3 TMl 1 1A1 Past Original P.T 10 10 11 13 10 10 Limits Past Orig. Design 13 8 12 14 6 6 lA2 Basis lA3 Past Orig. Design 68 45 64 70 33 30 Basis lA4 Future Plant Data 24 29 45 43 51 32 1A5 Future Plant Data 125 148 228 220 260 162 EVENT TOTAL 240 240 360 1360 360 240 Note: 1) Pressurizer level versus time is identical for all five transients.

2) An appropriate number of heatups is included in Transient lAl to account for hot functional tests.

O 3) This table is based on data as of November, 1990 and does not V

include the ANO 1 heatup in late December,1990.

4 31 l

l

1 Table 4-3. Events Affecting Surge Line Flow for Plant Heatup and Cooldown O Pressurizer pressurization (affects letdown rate and RCS volume and mass) ,

Purging of Pressurizer nitrogen along with RCS_ fill / vent operations Degassing Pressurizer using heaters / spray Adjusting Pressurizer level setpoint Controlling Pressurizer level in auto (w/ valve and controller deadbands) l Drawing of Pressurizer chemistry sample Forcing fill of RC loops by pressurizing Pressurizer PORY testing Control Rod Drive and RCS venting Decay Heat Removal System testing (i.e., throttling DHR flow)

Starting makeup pump Adjusting of RCP seal injection / return flows (adjusting seal backpressure)

Adjusting flow in RC pump seal bypass lin '

HPI check valve testing RCS heatup without RCP operating (throttle back on LPI cooler flow)

Closing letdown orifice manual bypass Drawing of vacuum in steam generator (OTSG) .

Starting first RC pump with thermally stratified RCS Starting first RC pumps Cycle auxiliary spray valve (for flushing of pressurizer)

Venting of Pressurizer to Letdown Storage Tank (LDST)

Changing RC pump combinations (affects minimum bypass spray flow)

Adjusting letdown flow rate with increasing RCS pressure Actuating spray (spray controls in hand)

Opening turbine bypass valves Testing Core Flood systen.

Adjusting heatup rate with turbine bypass valves (TBVs) (heatup hold / proceed)

Pressurizer spray controlling in auto, heaters on for boron equilibration Opening spray line block valve

\djusting makeup nozzle warming flow (s) (nozzle-warming throttle valves)

Placing TBY's in auto, steam pressure maintained within control band Adjusting boron concentration in RCS (changing MU/LD)

Incres. sing power to 8 percent Surveillance testing Degassifying Pressurizer and RCS (vent pressurizer to waste gas header)

Placing TBV's in hand and shutting down turbine Tripping turbine 4 32 4

--_.,.__n _ . - . - _ _ - _ - _ _ . _ _ _ _ _ - -

Reducing core power and reducing Tavg to 532F Sampling boron concentration in RCS ,

Adjusting TBV positions for desired cooling rate (TBV's in manual)

Performing Safety injection functional test Controlling Pressurizer level in auto (w/ valve and controller deadbands)

Discharging CFT's to RCS (test at <700 psig)

Closing HPI injection and bypass injection valves (at <350F)

Securing steam generator hot blowdown Initiating hot soak of SG's (optional)

Decreasing pressurizer level control setpoint to 100 inches Valving in the LPI system and starting 1 LPI pump Throttling LPI cooler flow to establish proper RCS cooldown rate Securing OTSG's by closing TBV's Filling OTSG's for Fill, Drain, Layup (optional)

Lining up and using auxiliary spray Adjusting LPI cooler outlet temperature to cold leg temperature Securing last RCP (RC loop temperature differential increases)

Adjusting LPI cooler outlet temperature to resume RCS cooling Securing RCP seal injection / return flows Opening letdown orifice manual bypass Adjusting letdown flow to lower pressurizer level to about 115 inches Regulating LPI flow as RCS is depressurized Going to 1 LPI pump Raising and lowering pressurizer level when on DHR Cooling loops by fill, soak, drain of SG (optional) l 4-33 l

L .

- Figure 41. Plant Data taken at Oconee Unit 1 l

. during HPl Check Valve Test (2/10/89)

(see Appendix B for details on the events occuring during this period)-

i i 135 400 - NOTE: Temperatures takenin fi Top Temperature surge line lower hortrontal section; ij 3

= 5 feet from the pressurt2ef.

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0 O 20 40 60 Time, hours -

Tpzr Tpzr Tres Tror s 1A4 1A5 -1A4 1AS-

. 4-43

~

i O '

Figure 4-11. Development of Heatup Design Bases Transients l i

Plant Operatkmal Plant Opek'ag Discussion with P-T Utitty Records- Ongsnal Desegn Data Procedwes Plant Operators Limits Event Count Bases __

l i 1 _ _ _ . . _ _ . _ _ _ _ . _ _ _ _ . . _ _ _ _ . _ _ _ . ,

Design Fases Generic SW6;cn of Surge Line Boundary Conditions Future Operaton_

Representog Past Opmahons (Transients 1 A1.1 A2,1 A3) g

. .I. .. +

j J Generic Specification of S>rge Line B ' %wy Conditions f __

Representing Future Operations (Trt;. - s 1A4 and 1 AS)

Y _ _ _ _ _ _ . __. __I _ _ _ _

Bomsary Condmonsinput to l Thennes Stratincation/ Striping Correistions

. t

-t 1

1 y _.

j New Surge Line Design Bases Transients i

i i

J -

se

O, Figure 4-12.

O P1 ant Heatup Transient IRI O ,

Temperatures - Pzr, Surge Line Ctop/ bottom), RC 700 l }'

/. ) i . I II T ggg. b  !!

l 8 '

~1 i

)kl /' ~~

f 8

! 1.! t l

_s . ._

f \

• • )..~ . - --_ . . b .-

,/ I f

\[\

500- -

T , , ,

7 v,I 1

^ 400 I 1

j D

b \ lI b l '

h '\ ( f o_ 300 s .i, V.,

x

\n f \eh

1 200

/ ( ;1 f

l

/ ,/

\' \ \ l

) 1

\

/ LF_h N ,

Top

( ." - - Bottom rpzr

_ - _ _ _ _ . Trc 0 80 20 40 GO 0

Time Chr)

z. ~. .. ,

O m

U <

Figure 4-13. Determining the Number of Events for Each Heatup Transient Past Operation Future Operation

& e2

, Heemfvmits be*Oyen m-I -

.( (

Time Ratio TemiNumber Totsf Number HFT to d Huetups the d Everts Hestup Events First 2 years u

't .( U Number d Equiv Number Number d - Number d 7

Future of Hestups Worst Case + Romeining Hhst.

Hestup Events for the HFT Transients Hestup Events

?

8 l

i' 1' Events without Events with Events wittout Events with Doesy Heat Decoy Hoot ~

Decay Heat Decay Heat (15) (5M)

- (1M)

(5M) .

p u

-U U p 1r Number of Nur9ber of Number of Number of Number d Transient 1A4 Trenolert 1A5 Transient 1A2 Transient 1A3 Transient 1A1 Events Events Events Events Everts 9

e 4

f

.n.

l ,

i i

i .

L- -i l Figure 4-14. Cooldown Design Bases Transient Temperatures  ;

700 .

- 1 600 -

t .

s Tpzr.1B1 s n S 500

. \ \

• \

' . . . . . . . . . . . . . .. ... .. .Tptr

. . . . .182

( ' ................................,

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=

1 i-

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g .-

t 300 - \ .

\

Tres .181.182 -.,

\. ... . ......._.... ...

200 - . --

., -l s ,g - i

.100 0 20 40- , 60 l

Eme, hours -

Tpzr Tpzr Trcs . >

-1B' 182 .1B1,1B2 :

a 1

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4-47 ,

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- -e - a ,. ,:. , - + , , , , ~ . , - , , ,~,<w.. ~

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q U

O .

.5. PIPING l ANALYSIS-s 5.1. Structural ~ Loadina Analvsis ,

The purpose of the structural: loading analysis it to generate-the internal forces 1

and moments in thefsurge line for the thermal-stratification conditions defined:

i in the. design basis transients.

'5.1.1. -Choice of Comouter Proaram The Structuralit.oading _ Analysis'of the surge:line- has been performed using th'e:

computer. program ANSYSl(reference 17). The decision 'to use ANSYS is based on the fact that both the straight pipe elevant STIF20 and the pipe elbow element STIF60 have been used extensively in the past. In addition, these two'ANSYS elements =

allou the user to define a variation.of. temperature onLthe pipe cross-section.

This capability is necessary for thermal stratification ~,- as a top-to bottom O _

-temperature difference exists in=:the horizontal = portion of the; surge-line.

1 5.1.2. Buildina of Mathematical.,Hgdgl To accurately simulate the:o. formation of the surge line as a result of thermal  ;

itratification, an " extended" mathematica1Leodel has been built. It consists of the pressurizer, surge line, hot leg, reactor vessel,- and. steam generator. .The boundaries of this " extended" mathematical model are.the pressurizer support at-

~

.'approximately mid-elevation of- the pressurizer? (where a stiffness: matrix is.

applied .to--simulate the flexibility of the support? frame) . the= basec of the reactor vessel skirt, .and' the base - of - the . steam generator pedestal. The

. mathematical model--of the surge line is given in Figure 5-1. The purposeLof" adding _ the pressurizer at one: end Lof the surge-line and the hot-leg,. reactor -

vessel, and steam generator at the other end is twofold:- it allows;the surge line nozzles to experience the correct-displacements and rotations due to the global tt.erma1= expansion of the primary system, and it correctly- simulates the ,

h- 5-1 2___._________ ._aa___-_a.___ - . _ _

l l j flexibility of the comp'onents adjacent to the surge line (the pressurizer and the-  !

hot. leg).

5.1.3. Non-Linear Temnerature ProfileL ,

-7 As mentioned in Section'5.1.1,= both.the straight pipe element;and the pipe elbowf (

element allow the user to define a variation.of temperature on the.' pipe. cross .  ;

~

section, the maximum temperature being at the top:of the' pipe,;and the minimum-temperature being at the bottom. of. the pipe. This variation of temperature.-

however, can only be applied as:a-linear function on the pipe cross-sectionL This is a. limitation,.as the~ temperature' measurements taken.during heatups~at L

Oconee Unit 1, in Februar: 1989 and again in' October 1989, have shown thatlthe: -

temperature profile on the pipe cross-section is non linear.' i To study the influence of a non linear temperature profile on the pipe cross -

section, the concept of " equivalent:11near temperature profile" is introduced: .

the equivalent linear. temperature profile is;the linear temperature profile which .

generates exactly: the same rotation of theE pipe cross-section iasj the one generated by the non-linear temperature ; profile 1 < The main steps of -thatinon- ,

linear-temperature profile study.are-as follows

L .

O . ivaluation of the non-><near tem,erature orofiies measured at Oconee Unit 1-(outside surface temperature profiles). ,

Selection - of the' most ' critical of those< non-linear temperature profiles (temperature profiles which experience a constant' maximum temperature for a considerable height-at' the top of the pipe cross = ,

-section, and a constant?minimu's temperature . for aE considerable' height at the -bottom of the pipe).

.- Selection of a simple temperature' profile at the-insidelsurface of- 1 the pipe' cross-section.

Finite element conduction run;to, arrive at the' temperature profile-I at'the outside' surface of the pipe cross-section.1

. Repetition of the previous two steps.:in an iterative manner, in an.

effort-to determine the temperature profile at3 the inside surface

.which - matches ' ' almost -. exactly - the'. "most critical" non-linear temperature profile atLthe outside. surface.

5  ;

i 4-. ~~ . - m. _ . . , , , _ ., .......,~m,_ , .. .s .w .,..,A. -y ....#- _,.,,-- .m.M i, , v.n w ..r , .- u_y

These iterative analyses have been perfomed for two non-linear measured temperature profiles at the outside surface: one for_ the Oconee Unit 1 February 1989 heatup and the other for the Oconee Unit 1 October 1989 heatup. These i measured temperature profiles have been chosen for their particularly non-linear; profile.

From the results of finite element conduction runs for which the calculated-outside temperature profile corresponds almost exactly to the measured outside temperature profile, the average temperature profile can be calculated easily.

This average temperature profile is then applied as input in the mathematical-formulation of the " equivalent linear" temperature profile. In addition to being a function of the average temperature profile, this mathematical formulation. is also a ft.nction of the modulus of elasticity and the coefficient of thermal _

expansion, which are both degn~ dent on the temperature.

The non-linearity coefficient is defined as the ratio between the top-to-bottom temperature difference of the equivalent-linear temperature profile and the' top-to-bottom temperature difference of the measured non-linear temperatura profile.

As the analysis described above has been performed for two non-linear measured temperature profiles, two different non-linearity ' coefficients ere calculated:

1.32 and 1.35. It is not surprising that these two coefficients are not equal, as they are, in general, a function of the following three parameters: the temperature at the top of the pipe, the temperature at the bottom of the pipe, and the elevation of the fluid interface centerline, the fluid interface being the height on which there is a non-negligible variation of the temperature at the inside surface of the pipe cross-section.

Finally, the mathematical formula of the non-linearity coefficient has 3een developed, as a f metion of the temperature at the cop of the pipe, -the temperature at the bottom of the pipe and the elevation of the fluid interface centerline. The variations with temperature of the modulus of elasticity and of the coefficient of thermal expansion have been embedded in that-mathematical formula.

O .

s-3

-5.1.4. -Verification Run for Disolacements During i the Oconee Unit l' February 1989 - heat-up, the _ temperature and the displacements of the surge line have been recorded.

< During that heat-up, the' maximum recorded top-to-bottom temperature difference  :

is'280F (T ,- 403F; Tm = 123F). For that exact time point, the temperature .t at other locations of the surge line and the surge line displacements are known'..

At that particular time point, the top-to-bottom temperature difference varies along the horizontal portion of the surge line from 280F maximum to 200F minimum.

Also, the top-to-bottom temperature profile is not exactly the same -at :each location where temperature data have been measured. Therefore, the non-linearity-coefficient is calculated at each one of these locations, using the mathematit 41~

formula developed as explained in Section 5.1.3.

The resulting equivalent linear temperature profiles have been given as input for the horizontal portion of the surge line, together with the other temperature i data measured at the same exact time' point in the surge:line riser, in the hot-leg, and in the pressurizer.

Using the " extended" mathematical mode'l _of the surge line, (documented in Section-

'O- 5.1.2),-an ANSYS computer run has been performed. ~ The resulting displacements:

have been compared with the 'measur_ed displacements, as shown in Figures 5 2 ~

~

through 5-4. These. comparisons have shown a very good agreement between- the.

calculated and the measured displacements. (See Figures 5-2 through 5-4.) ,

This has verified that the " extended" mathematical model of the _ surge line is'

, accurate for the generation of the internal forces and moments in the structural

. loading analysis. It has' also verified the method used to correct for the influence of the non-linearity of the temperature profile ~ on - the pipe . cross-

=section.- (An ANSYS - computer-run without the non-linearity = coefficient gave displacements that do_ not compare well with the measured-displacements.)

5 .1.' 5 . Structural Loadina Analysis for the Thermal Stratification Conditions-The thermal . stratification conditions have been defined' in the design basis transients documented in Section.4.5. In these design basis transients,- the thermal-stratification conditions are evenly divided in peaks and valleys, where the peaks are characterized by maximum top-to-bottom temperature differences and i

5-4

" ~ ~ ..,e,, --n,m-~---. v. . -e an... .,ro- nn-w+~ w . , , --un-s-.e si

- the' valleys by minimum topito botton_ temperature . differences. _ Among the peak -

conditions, the three1most critical . thermal stratification conditions occur-

~

during the heatup. transient IA1. These three thermal stratification conditions are characterized by top-to bottom temperature differences of 397Fi-393F and ~

- 386F. <

' For each one- of these three stratification conditions, a structuraltloading -

analysis has been performed,'us'ingLthe ANSYS mathematical model documented in 3 Section~5.1.2. For each of these stratification' conditions,_the topito-bottom temperature difference has been multiplied by the non'-linearity coefficient, to '!

arrive at the " equivalent linear" top-to-bottom' temperature difference which.has- -

~

been given as'an input.for the ANSYS computer run. Also. the pressurizer and hot-leg' temperatures have been defined in the designb' asis traniients and are given =

as input.. The temperature in the_ riser of:the surge line;is given as T , in the>

lower half of the riser.(the temperature at the' top of the pipe in the horizontal. ,

run), and as the average between T,, and Tw, gin the upper half of the riser, r This temperature distribution .in the riser -is known . from thez measurements 3

performed Lat Oconee Unit 1, and is the,same:as_ already used in the Verification [

Run-of Section 5.1.4. -

In : addition !to the three- ANSYS computer-runs 7 performed for. thesthree most critical thermal stratification. conditions, the structural loading. analysis has

j. also ; been performed for a spectrum eof sevent other thermal 'stratificatien .(

a conditinns defined in the design: basis-transients.";These conditions have been chosen to cover a wide range of different top-to) bottom temperature differences.- J Also,- the structural' loading. analysis > has been7 performed. for- the 100% power thermal conditions characterized by.a complets lack of top-to-bottom temperature-difference.

l As a result of each thermal-stratification condition above, the internal forces:

and moments in the surge line have been calculated (in" separate' ANSYS computer- ,

runs). Furthermore, having these 11 sets of internal forces and moments allows i the generation of an interpolation scheme to determine the-internal forces and-I moments everywhere in the: surge line. -The o variables: included in- this interpolationischeme are the locationLon' the Lsurge^11ne (coordinates),- the 1

' " equivalent linear" top-to bottom temperatureLdifference, the . average _ between T,, and Tw,,,, and the temperatures 'of thel pressurizer and the hot-leg.

5-5 5

a- - - =

• g - , ,,,---=we w, ,,,e sommre ce , w e , w e, e, wv . v e -, e w -- ww ev-, ow r g w rw, i

1 4

5.2. 2 Generic Stress Indices for the Surae Line' Elbows :

The thermal expansion secondary stress has tI bef verified-in NB 3600 through thef Equation - 12 ' stress 'which11s: required- to _be within the < 3*$m allowable. This: .

Equation-12 ctress contains the stress"index C,. For the elbows, N8 3600_ gives; j

-the same valuestfor the secondary stress index C,~ and for K,

• C,, which is the--

1 multiplication offstress indicesitolbe usedlfor the' elbow peak stress:(K, is; _

1 defined as 1.0 for. the: elbows). Using;the definitions 'of: secondary and- peak stresses given in:NB-3213.11;and NB 3213'13,. . the C, and K, stress: indices have. j been- recalculated > for?the surge - line elbows..o The followingi finite; element-1 analyses are for all four surge line elbows, as-theyf are identical to each_other

__ a I

(90 ' degree elbowsJ10 inch schedule 140[ stainle s steel A' 403, grade WP316). _

a The computer program ABAQUS (reference 19) isiused for the purely Leitstic!and :

elasto plastic finite element analyses described in the next subsections.c ABAQUS:-

is a general-purpose finite element program for li_near an'd non-linear structural

applications. ABAQUS has been chosen for these finite element analyses, because, whila it is well suited for purely linear analyses,-the overall" program desig'n; is dominated by features especially suited for Lnon-linear l and elasto-plastic-problems.

5.2.1. Prelv- Elastic: Finita' Elamant Analvsis ' -

A finite element model'of thel surge line elbow has:been builtLas shown in Figure.

5 5. A purely elsstic- finite element :analysischas ben.' performed, using the:

computer program ABAQUS,- for the following two-loading conditions:' . in plane bending of the elbow,: and-out-of-plane bending.of the elbow'.

The most critical stresses.from these two loading cases are compared _ with the-stresses which'can'be calculated by using.the formulas given;in Table;NB-3685.1- -

2. From both loading conditions; the highest stress index for?these maximum-stresses occurs when applying in-plane bending: : 2'.32Lus'ing Table -NB 3685.1-2, and 2.33 from the ABAQUS ' finite - element- Analysis._ = This -verifies' the- ABAQUS finite element model and-reinforces:the fact that the formulas given in Table NB-

'3685.1-2 are calculating peak stresses. . -

5-6 i

_ . . . .- e , -

. . . - - -e _ - . _ . - . - - - - - . - . _ - . - - _ _ - .- - _ - - .

1 5.2.2; Elasto-Plastic Finite El nt? Analysis L

O t#47 th '#ri c r:th a

• tr th ei t r14 # r **

'elbow', an ABAQUS elasto-plastic finite element analysis' has been perfoj ra 'i - .

the1same two. loading conditions as--in subsection 5.2.1: - in-plane bending andf .

- out-of plane bending.

L In the case 'of in-plane bending,ithe. highest stress'in the elbow reaches the 3*Sm

-l allowable at the loading level l for which K2

• C2 is' equal'to 2.33. JAfter_ passing - I this' loading level, the inside and outside fibersion:the' elbow thickness enter:

the elasto plastic domainLin very localized regions-of the cross-section and do i not contribute further tolthe strength.of the elbowf Asta result, the' surge line-  ;

^

elbow no longer behaves in a= purely _ elastic manner. : However,'wh'en plotting these -

displacements asia function of;the increasing loading;1evel :it can be seen that'-

the ~ generain distortion .of. the surge-line 'elbowiremains-1an: essentially linear 7

function of the loading level. .This; meansj that,ifor aEreasonableL r_angelof?

~

loading above the loading level which causes(initialiytelding,4thelhighAst} '

e stresses at the- inside and 'outsideL fibers on the7 elbow thicknesst do noti

contribute-. significant1y' to ' the distortion ; of { thei elbow. .-(This

ist becausei  ;

initial yielding is localized.)  ;

O .

In other words,_these' stresses are peak; stresses-which should be considered'in; i

the: fatiguez analysis, but not- secondary ' stresses, Las1they; do not causeEany:

appreciable > distortion. lIn thel case of out'-of-plan'elbending,3 the1 surge line; i

' elbow behaves-in a very similar fashion as inzthe casecof in-plane bending.

5.2.3. Calculation of a Generic Stress Inder C2 for~ Sacondarv' Stress in the- ]

Surae Line Elbow The determination of.the~ ger.eric' stress lindex Cpfor. calculating the maximum secondaryJstress in the surge lineielbow has been based on;the fact that the-secondary stress is the stress which contributes to the~ general; distortion of the; elbow. Therefore,1 the maximum state-of-stress-has been linearized'on theTpipe i cross-section, and;the resulting stress at the outside fiber of the pipe cross - '

.section is the maximum secondary stress, after an additional factor of safety is -

applied (using the fact that this'is a displa' cement controlled phenomenon).: j v

5-7 n~ sum-., e,w e,em o w m-, w e,- ,.-rw- -, e r m + - .r v ee-ev e n +

The resulting stress indices C, from the finite element analysis are as follows:

/

() . For in-plane bending, C, - 1.30

. For out of-plane bending, C, - 1.58 Therefore, the generic stress indices for the surge line elbows are: C, - 1. 58 and K 2

- 1.48, leading to a total peak stress index Kr

• Cr equal to 2.34. These stress indices have been used in the Fatigue Analysis of the surge line.

5.3. Verification of NB-3600 Eauations (Ecuations 12 and 13. and Thermal Stress Ratchetinal When accounting for the most critical thermal stratification' cycles, the Primary Plus Secondary Stress Intensity Range, Equation 10 of NB-3653 (reference 18),

exceeds the 3*Sm limit. This occurs typically for the thermal stratification cycles associated with the very high top to bottom temperature differences in the surge line (such as the past heatup transients). For all other thermal stratification cycles, the Equation 10 Stress Intensity Range is verified to be less than the 3*Sm limit. To be able to perform the simplified elasto-plastic fatigue analysis in accordance with NB-3600, it is necessary to verify Equations 12 and 13 of NB-3653.6, and the Thermal Stress Ratcheting Equation of NB-3653.7.

These verifications (described below) are performed using the 3*Sm limit based

(] on the ASME Code allowable yield stress for the material.

Equation 13 consists of calculating the primary plus secondary membrane plus bending stress intensity, excluding thermal expansion, and comparing the total resulting stress with the 3*Sm limit. Equation 13 stress is due to dead weight, l operating pressure and Operating Basis Earthquake (OBE). The surge line does not l

contain any materi;l or thickness discontinuity. Therefore, the third term of l Equation 13 stress is equal to zero (no variation of modulus of elasticity and no abrupt variation of average temperature in the axial direction of the piping).

Equation 13 stress has been shown to be acceptable at every surge line location Fr all six B&W loered-loop plants. The maximum Equation 13 stress occurs in the elbow just belou the pressurizer (elbow D) for the TMI Unit 1 surge line.

It is equal to 26.8 Ksi, which is 56% of the 3*Sm limit of 48 Ksi.

The verification of Thermal Stress Ratcheting consists of comparing the highest occurring Delta T, range with an allowable value to be calculated in accordance with NB-3653.7 of reference 18 (Delta T, range is the range of the linear b) a 5-8

through wall temp:rature gradients). This verification is performed in the fatigue analysis of the surge line described in subsection 5.5. This is because the right hand side of the equation (the allowable Delta T range)includesthe 3

yield strength value which has to be taken at the average fluid temperature of the_ transient under consideration. However, for a yield strength value taken at the mhximum possible temperature in the surge line (650F), the allowable Delta  ;

T range is still more than 15% higher than the maximum Delta T3 raage occurring 3

during the thermal stratification cycles. Therefore, Thermal Stress Ratcheting is verified to be acceptable (by more than 15%).

Equation 12 consists of calculating the secondary stress range due to thermal expansion and comparing it with the 3

• Sm allowable. To do so, the internal forces and moments due to each one of the three most severe thermal stratification conditions have been compared.

Applying the thermally adjusted internal forces and moments from the most severe thermal stratification conditions, Equation 12 secondary stress has been verified at every surge line location, using the simplified ASME-Code equations and the conservative generic elbow stress indices (derived earlier), except for three elbows. These three elbows are elbows B, C and D (see Figure 3-1).

The three elbows which failed the simplified Equation 12 have been addressed via the elasto-plastic finite element analysis. For each one of these elbows, the thermally adjusted internal forces and moments from the most severe thermal stratification conditions have been applied on the elbow finite element model documented in Section 5.2. The method used to calculate the resulting maximum secondary stress was exactly the same as the-one used for the generation of the generic stress index C, in Section 5.2.4. For these elasto-plastic finite element analyses, the 3

• Sm allowable is taken (conservatively) at the maximum possible temperature in each elbow. This maximum possible temperature is, for all three elbows B, C and D, the temperature at the top of the horizontal portion of the surge line. The maximum resulting secondary stress is calculated to be smaller than the 3

• Sm allowable.

5-9

5.4. Develooment of Peak Strg1 31 5.4.1. Peak Stresser due to Fluid Flow S

The design basis transients specify the temperature data necessary for the Structural Loading Analysis documented above (subsection 5.1.5). In addition, they specify the maximum flow rate in the surge line for each time span between two consecutive PV's (either between a valley and a peak, or between a peak and a valley). Also, the maximum variation of the fluid temperature with time can be found in the design basis transients for the nozzles.

During an outsurge, the hot temperature fluid from the pressurizer flows on top of the relatively stagnant lower _ temperature fluid. This leads to a hot temperature flow,-into a relatively cold inside wall. The maximum flow rate for the transient has been' assumed to be constant for conservatism; and the variation of the fluid temperature has been taken at the highest temperature gradient given in the design basis transients for all the outsurges. The flow cross-section has been reduced to the upper portion of the inside pipe cross section, thus maximizing the flow velocity and the through-wall temperature gradients Delta T3 and Delta T,. These temperature gradients occur at the top of the pipe.

During an insurge, the cold temperature fluid from-the hot leg flows through the surge line. As a result, the temperature at the bottom of the pipe does not change significantly, but the top of the pipe experiences a temperature decreate from Tp (top, peak) to T y (top, valley). Therefore, the insurge is analyzed similarly to the way the outsurge is analyzed: cold temperature flow, into a relatively hot inside wall. The maximum flow rate for the transient has been assumed to be constant for conservatism: and the variation of -the fluid temperature has been taken at the highest temperature gradient given in the design- basis transients for all the insurges. The flow cross-section is the whole inside cross-section of the pipe. (Buoyancy is minimal since hot leg water has approximately the same temperature as the water at the bottom of the pipe.)

The maximum through-wall temperature gradient s Delta T and Deltaa l occur at the 3

top of the pipe.

The Peak Stress -Intensity Range is a direct function of the through-wall temperature gradients Delta Ti and Delta T2 . (NB-3600 Equation 11). Both the A

-V 5-10 1

insurges and the outsurges have been utilized to obtain the total Peak Stress Intensity Range.

The through wall temperature gradients - Delta (3 and Delta 'Tf calculated as described above are very conservative._ This is due to the fact that, for both the outsurges and the insurges, all temperature data,_ flow rates and fluid temperature variations are taken at their conservative extreme values (either maximum or minimum). Also, no consideration has been given to the loss of temperature as a result of thermal conduction around the pipe cross-section (which would mitigate the effect of the extreme fluid temperature ramp assumed).

).4.2. Peak Stresses due to Thermal Strioino The thermal striping temperature data are given in the design basis transients.

These temperature data are maximum inside wall temperature ranges occurring in the region of the fluid interface centerline.

! To calculate the temperature distribution through the thickness of the surge line at different points in time, a mathematical model -has -been built, using the Finite Element program ANSYS. The time-dependent wall temperature has been simulated as a " cut-sawtooth" wave which is very close to a sine wave. From the O' experimental thermal stripino erk described in Section 4.3, it is known that the thermal striping fluctuations ha <e a pertod. of approximately l'.0 second.

Therefore, a sensitivity analysis is performed to _ determine for which ' thermal striping period the peak stress intensity range is the highest. The tnermal l

striping periods considered in this sensitivity analysis are 0.5, 1.0, 2.0 and 4.0 seconds, to cover the entire rar.ce of periods which- could be expected.

For each thermal striping period studied, the ANSYS computer-run produces time-dependent temperature profiles through the pipe thickness. There are two extreme temperature profiles: one above the average temperature and with the inside wall temperature reaching a maximum, and the other below the average temperature and-l with the inside wall temperature reaching a minimum. For:these two temperature l profiles, the linear and non-linear through-wall temperature gradients have been calculated, leading to the maximum peak stress intensity range.

This evaluation has been repeated for each one of the four thermal striping periods. It is found that there is not one most critical thermal striping period.

This is because the maximum peak stress intensity range is a function of the 5-11

_- _ __. - . . _ _ _ _ -~_ __ _ _ _ . _ - _ . . . _ .

stress index K3 which acts on the peak stress contribution from the linear g

V through wall temperature gradient Delta T , but not on the one from the non-3 linear through wall temperature gradient Delta T g(which does not have any stress index). For a stress index K3 equal to 1.7 (girth butt welds, as-welded), the most critical period is 4.0 seconds, because.the higher the period, the greater the linear temperature gradient Delta Ti is. On the other hand, for K3 equal to 1.0 (locations remote from welds), the most critical period is 0.5 seconds. In that case, the maximum peak stress is the so-called " skin stress", due to Delta T

2 (in the surge line,1.7 and 1.0 are the only two possible values for K3).

For the Fatigue Analysis of the surge line, the peak stress intensity range due to thermal striping has been taken at its maximum possible value: at the as-welded locations, it is the peak stress intensity range corresponding to a striping period of 4.0 seconds; and at the locations remote from welds, it is the peak stress intensity range corresponding to a striping period of 0.5 second.

5.4.3. Peak Stresses Due ti ,,.e Non-Linearity of the Temperature Profile As already mentioned in subsection 5.1.3, the top-to-bottom temperature profile on the pipe cross-section is not linear. In the Structural Loading Analysis of .

l l () the surge line, the " equivalent linear" temperature profile has been applied for the generation of the internal forces and moments. The difference between the actual non-linear and the " equivalent linear" temperature profiles is a relatively small temperature difference which is referred to as the Delts T 4 temperature difference, it causes a peak stress very similar to the one produced l by the non-linear through wall temperature gradient Delta T2.

The two most severe measured top-to-bottom temperature profiles have been analyzed (see subsection 5.1.3), these two measured top-to-bottom temperature profiles have been selected for their particularly non-linear shapes. Therefore, these non-linear shapes will lead to the highest possible temperature differences Delta T4 . To better study the influence of Delta T4 on the localized state of.

stress in the pipe, an ABAQUS Finite Element model has been built (reference 19);

and the temperature differences have been applied as input. The ABAQUS Finite Ehment analysis is performed for each one of the two measured top-to-bottom temperature profiles.

/

V 5-12

From the two ABAQUS Finite Element analyses, it has been found that the maximum peak stress intensity occurs at the inside radius of the pipe cross-section. This maximum peak stress intensity takes into account all localized effects in the pipe cross section, including the localized bending on the pipe thickness.

A correlation has been developed to calculate the maximum temperature difference Delta T , as a function of the top-to-bottom temperature difference and the 4

elevation of the fluid interface centerline, and give the maximum peak stress intensity in the pipe cross section, as a- function of the maximum temperature difference Delta T4, the top to-bottom temperature difference and the elevation of the fluid interface centerline.

5.5. Fatique Analysis of the Surae Line.

As described in Section 5.3, Equations 12 and 13 of-NB-3653.6 and the Thermal Stres:; Ratcheting Equation of NB 3653.7 have been verified to be acceptable.

Therefore, the fatigue analysis of the surge line can be performed-in accordance withNB-3653.6(c). The verification of: Equations 12 and 13, the Thermal Stress Ratcheting verification, and the fatigue analysis of the_ surge _line have been performed in accordance with the 1986 Edition of NB-3600. This is a requirement O of NRC Bulletin 88-11: " Fatigue Analysis should be performed in accordance with the latest ASME Section 111 requirements incorporating high cycle fatigue" (reference 1). A Code reconciliation has been performed to remove all concerns associated with using the later Code (1986 Edition) for this fatigue stress analysis. For this Code reconciliation, each plant's surge line stress report has been reviewed (the code of record for piping analysis is not the same for all the surge lines). The Code reconciliation has shown that:

1) research performed since the publication of the B-31.7 USA Standards (1968 and 1969) has led, in general, to more sophisticated mathematical formulas for the stress indices,

2) the allowables used in this final analysis are either equal to or smaller than the ones used in the original stress reports (therefore, the allowables used in this final analysis are conservative),

3) the fatigue analysis of the surge line incorporates the fatigue curves up to 10" cycles, in comparison with the fatigue curves of the B-31.7 USA Standards (which are restricted- to- 106 cycles, and therefore do- not 5-13

I l

i contribute any additional fatigue usage for relatively low alternating  !

stresses).

For each of the thermal stratification peaks and valleys from the design basis transients, the total Equation 11 peak stress intensity has been calculated. The complete " loading map" for the peak stress intensity values is as follows:

. Moment loadings due to thermal stratification (" Equivalent linear" top-to-bottomtemperaturedifference_),

. Internal pressure in the surge line,

. The additional localized peak stress due to the non-linearity of the top to-bottom temperature profile, and-

. The maximum stress (for a given thermal stratification condition) between the peak stress due to thermal striping and the one due to Fluid Flew. The maximum is used, because. these two peak _ stress intensities occur at two different locations in the pipe cross-section during the same thermal stratification condition (in the region around the fluid interface centerline for thermal striping >

and at the top of the pipe for the Fluid Flow conditions). -This maximum peak stress intensity is- added for the thermal stratification peaks and subtracted for the thermal stratification .

l valleys.

A sort of all the total peak stress intensity values has been performed, and a selection table nas been built for the combination of the thermal stratification peaks and valleys into pairs. This selection' table is. built in such a way that the peak stress' intensity ranges are maximized.

l l For each pair of thermal stratification conditions (one peak and one valley), the alternating stress intensity (Sa) has been calculated, as a function of the peak stress intensity range and of Equation 10 primary plus secondary stress intensity range. The usage-factor associated with the Sa value has been calculated,-using the " extended" fatigue curves of the ASME-Code, Section III, Appendix I (up to 10" cycles). A summation of all usage factors on all pairs of 1.hermal stratification conditions has resulted in the main' fatigue usage. This main fatigue usage includes all the thermal stratification conditions characterized 5-14

., _ g , _ _ , , ,

l n by a top to-bottom temperature difference. Also, the surge line internal moments-V due to an Operating Basis Earthquake have been added conservatively to the most critical-thermal stratification-internal moments, for consideration of future OBE occurrences.

As required by ASME-Code NP ;600, all stress intensity ranges from the different - 4 contributions given in the ioading map".above have been combined to generate the maximum total stress intensity range to be used for fatigue.

Also, as part of this fatigue analysis, the- Thermal Stress Ratchet has oeen-verified to be acceptable, in accordance.with NB-3653.7. This verification is performed for every range of linear through wall temperature gradients (Delta T3 range).

The following fatigue contributions have been added to the main fatigue usage:

. Fatigue due to highly cyclic thermal striping _ ranges. The design basis transients define, for each event, the number of occurrences-of each thermal striping range.

. Fatiguo due to the-future Operating Basis Earthquake ranges-(cycles which are not associated with the thermal stratification events).

. Fatigue due to additional Fluid -Flow conditions which are not associated with- significant thermal stratification conditions.

The fatigue usage factor due to the highly cyclic thermal striping ranges has been found to be in the following ranges: between 0.10 and 0.15 at the locations with a stress index K3 equal to 1.7 (girth butt welds, as welded), and between 0.06 and 0.09 at the locations with a stress index K3 equal to-1.0 (locations remote from welds). The reason for the 50% difference between the extreme values in each range is the fact the number of design transients is not the sa'me for each B&W lowered-loop plant.

The fatigue- usage factors associated with the future OBE- cycles and with the additional fluid flow conditions are -very small. However, in order to complete the total fatigue picture, they have been added to the sum of the usage factors from the main fatigue and from the highly cyclic thermal striping ranges, t

5-15

N 5.6. Fatiaue Analysis Results for the Surae Line Table 5-1 is a plant-specific table of the total fatigue usage factors for a 40 year plant life (including past and future fatigue). This table is plant-specific because the number of occurrences of the ever.ts defined in the design basis transients is unique to each B&W lowered-loop plant.

l t

O wJ l

i

\.) 5-16

- . . . = - . - - . - . . . . . - - . - . - . . - . - . _ . . - . - . .. . - .... - - . .. . .- ..

~

1 0 1 8,e e-i. snecir4c 'et 1 r t49"e us se r>cters rer the serae ee 1

Surge Line Oconee Oconee Oconee CR 3 ANO 1 - TMI-l Locations Unit 1 Unit 2 Unit 3 Most Critical 0.48 0.48 0.47 0.37 0.38 0.38 Straight Most Critical 0.79 0.82 0.76 0.64 0.66 0.65 Elbow Second Most Critical Elbow 0.76 0.79 0.75 0.60 0.60 0.61 Drain Nozzle 0.35 0.37 0,33 0.25 0.26 0.25 Branch r

(See Section 6 for the pressurizer surge nozzle and- the surge line / hot leg nozzle).

1o All total fatigue usage factors are smaller than their allowable of 1.0, The highest cumulative damage is 0.82 (this value is for Oconee Unit 2).

The most critical locations are the same for all six surge lines: the highest cumulative damage occurs in the vertical elbow at the bottom of the surge line riser (elbow B); and the second highest cumulative damage occurs in the vertical elbow just below the pressurizer (elbow D). Elbows B and D are indicated in Figures 3-1 and 5-1.

l l

l 5-17

, -y-w r 4 ,m -m* - ,-www- _ -., , ,- e,,. ,v.,, er..~n4-*--m,e - - - -mi- -<y--n - -we,

O x/

Figur e 5-1. Surge Line Ma themo tical Moclel ELBDW A 4 32 5 HDT LEG 6

elements 1 through 5 7 ,

Jt.8: X= - 5 5.0 ' 1y y = - 1.50

• 8 f= 24 8.656' X

9 Z/ -

elements N 10 through' 13 10 7

11

/ N ELBOW B 12 9.0' (X = + 53.0')

4 13 Drain Nozzle 5

(]s /,for N S S S - 3, 4, 5 and 9.

4e elements -

s g ,

g 23 through 26

\

N 19 NSSS PLANT g 3 DCONEE I (X= 41.0') 20 4 DCONEE 2 Drain Nozzle 5 TMI 1 for NSSS- 7 and 8. N 21 7 CR 3 8 ANO 1 22 23 9 OCONEE 3 ELBOV C 24 26 Pressurizer elements 29 through 32 7 27

/ 28 /

32 /

ELBOW D 31 Q 30 29

,~

V 5-18

I ,

O O O Figure 5-2.- Comparison of- Surge Line Displacements, X - direction

(Calculated vs. Measumd with DT-280F)

' O.5 0.4 -

Q

.E .

O O v

O.3 -

3 6

6 0.2 -

8 2

n.

.9 m o .o1--

8 '-

~g- O I

X

- 0.1 -

'- 0;

-0.2 i .i  : -- i

.i 120 30 , 40

'O- .10 Distance from PRZR-oxis (feet)

O measured +- calculated

? +-- r---- a _

O O O Figure 5-3. Comparison of Surge Line Displacements, Y - direction . l

~

(Calculated vs. Measured with DT-280F) _

0.5  ;

0.4 -

e 0.3 -

a O.2 -

9

'o .

0.1- -

O O- ,

O '

--0.1 ' -

E-o i

-E -0.2 -

O w

j -0.3 -

a- --0.4 -

2 }

m O-

,N c -0.5 - -

o .o-

'--0.6 -

E O l

-0.7 -

.g s

I --0.8 -

>- ~0.g -

1 D

-i

~ 1.1 - -

l -12 -

i i i i i a i i i i e i i O 4 8- 12 16 20 24 2R 32 *

i j Distance from PRZR-oxis'(feet) .

O measured +. calculated I

i 1

O O O Figure 5-4. Comparison of Surge Line Displacements, .Z - direction ,

(Colculated vs. Measured with DT-280F) 0.3 0

m 0.2 -

U r .

2 o

~

3 0.1 -

6 E

8 U o

o A -

a Y

ro b l C

0 y -0.1 -

h D D ,

i

" O

-0.2 -

-0.3 i  : i i .

i i i i i O 4 8 12 16 20 24 28 32 Distance from PRZR-oxis (feet)

O measured + calculated i

l l

-O Figure 5-5. Finite Element Model of Surge Line Elbow

.El E \

l A

w n b

, 12){ & h @Y-24 Elements Around the Pipe O v.

w t2% inches -

a

.f O

5-22 g- >- "yw' '-s e-

i I 1 o

t

' 6. N0ZZLE ANALYSES In addition to the structural analysis _ of the surge line described in Section 5, detailed stress analyses of the pressurizerland hot leg. nozzles have been -

j performed to_ demonstrate compliance with the : requirements .of the ASME Code,;

Section !!!. -The thermal and pressure parameters for each nozzle 'are described. 1 in the design basis ~ transients of- Section- 4.5. 'In addition, Lat.h nozzle is subjected to piping loads from the surge line itself. These ? 6 have been taken from the piping analysis described in Section 5_ and the resultirg stresses have been combined with those from pressure and' thermal loading;.

Detailed descriptions of the analyses of the pressurizer surge nozzle and of the 4 surge line to -hot leg . nozzle are contained inLthe following sections. The analysis of the surge drain nozzle is part of the, surge.line' structural analysis-described in Section 5 (Table 5-1 gives the total fatigue usage for thel surge

! drain nozzle of. each plant).

1

$ 6.1. Pressurizer Surae Nozzle The purpose of this section -is to describe the evaluation of.the. pressurizer surge nozzle. The analysis of the nozzle, safe end,'and safe end-to-pipe weld I has been performed using the finite element > method as implemented by-the "ANSYS" computer cr. . Reference 17. The loads used for evaluation were the thermal and pressure loads identified in the-design basis transients fore the surge line t' ' cation (see:Section 4.5). _The acceptance criteria for the' evaluation' a requirementr, for Class 1-components of the 'ASME B&PV code, Section III 3 ,

T. edition ~ with no - Addenda, reference- 18.- The safe -end,- nozzle, and  !

urizer head were evaluated using the requirements of NB-3200. The safe end-to-albow weld was evaluated to the requirements of NB-3600.

6-1 1

, . ,e ,, --u.

6.1.1. Geometry The pressurizer surge nozzle is shown in Figure 6-1 and consists of a carbon steel nozzle welded to the carbon steel lower head. Both the nozzle and pressurizer head are clad with stainless steel to prevent reactor coolant fluid from contacting the carbon steel base metal. The nozzle and head are also protected from thermal shock by a stainless steel thermal sleeve.

A stainless steel safe end is welded to the end of the surge nozzle and then welded to the stainless steel surge line piping in the field. The weld between the safe end and pipe is critical since it is in the "as-welded" conditan and requires special consideration (stress indices) in the stress evaluation.

6.1.2. Description of loadinas The loadings on the pressurizer surge nozzle consist of thermal gradients, internal pressure, and external piping loads. Since the' pressurizer surge nozzle is vertically oriented, there are no significant thermal stratification loads (ie. temperature gradients from one side of the nozzle to the other).

The thermal gradients are caused by.the various temperature swings (peaks and O valleys) associated with the in- and out- surges of fluid between the pressurizer and the RC pipe. The temperature swings are defined in the design basis for the surge line. The thermal gradients and stresses due to these temperature swings are determined using the ANSYS finite element code.

The RC system pressure is applied to the internal surface, of the nozzle, safe end, and head. The pressures are defined in -the design basis and the resulting stresses are determined using ANSYS.

There are significant external loads developed due to the. heating and cooling of the surge 'ine as well as the stratiff ration in the surge line. The external loads, forces and moments, for each peak and valley - are given in the documentation of the surge line analysis. The stresses due to these moments are calculated using classical solutions (e.g.- Mc/I).

6.1.3. Djscussion of Analysis

- An axisymmetric thermal and thermal stress analysis has l'een performed using the ANSYS finito element computer code. The transient thermal analysis consists of

. imposing time dependent boundary-conditions (bul'( fluid temperatures and heat 6-2 se

~

1 3 transfer coefficients) on the finite element model. Nodal temperatures from the thermal analysis have been stored on magnetic tape for each iteration (time step) of the transient. The ANSYS postprocessor " POST 26" has used the nodal temperatures to calculate Delta T's between various locations in the structure.

Tables of the Delta-T's versus time for each transient have been used to determine when the maximum and minimum stresses are likely to occur. The nodal temperatures for each critical time step are input to the ANSYS stress routine for the determination of stresses. The ANSYS postprocessor, " POST 11", has been used to linearize the stresses at critical sections of the structure. Pressure stresses at the critical sections have also been determined using ANSYS and POSTil. Stresses due to external loads are hand calcul.ted and combined with the thermal and pressure stresses for comparison to ASME code allowables.

Due to the large number of temperature swings (peaks and valleys) associated with the thermal stratification transients it was not practical to evaluate each peak and valley as an individual case. Instead, a few " base cases" were created to envelop the large number of identified peaks and valleys. The base cases were chosen to cover both up and down ramps, various starting temperatures, various A ramp rates, and various changes in temperature (Delta-T). The temperature V distribution and resulting thermal stresses for each of the base cases have been determined using the procedure described above. A sumary of the parameters used in the base cases and the resulting stresses is given in Table 6-1. The actual transier,ts identified in the design basis were reviewed and each peak and valley have been assigned a representative " base case". The selection of " base case" l was such that the star'.ing temperature, ramp rate, and Delta-T of the actual transient were bounded by those of the " base case". The linear and maximum l

i thermal stresses from the chosen " base case" have been used directly for combining with str9sses due to pressure and external loads.

Pressure stresses foc a base case with an internal pressure of 2200 psi have been determined using ANSYS. The pressure stresses for each peak and valley have been determined by multiplying the stresses from the base case by the ratio of the actual pressure for the peak or valley from the design basis to the pressure used in the pressure base case (2200 psi).

The external load stresses for each peak and valley 'have been determined by taking the moments from the surge line evaluation at the time nearest the peak O 6-3 4

h I

or valley and calculating the resulting stress using cla;sical solutions (e.g. ,

O Mc/I).

The thermal, pressure, and exter.al. load stresses have been multiplied by the appropriate stress indices (Table NB 3681(a)-1) or stress concentration factors and then combined for determination of maximum stress and fatigue usage. The results are given in Section 6.1.7.

6.1.4. List of Assumotions/ Inputs Usod in Analysis

1) Since the surge line elbow is a 3D structure it cannot be included in an l

axisymmetric model, therefore it has been modeled as a straight pipe.

This change should have a negligible effect on the resulting stresses in the safe end to elbow weld. The cctual stresses in the elbow have been considered in the evaluation of the surge line, Section 5.0.

2) Because the exact fluid boundkry conditions on the inside surface of the lowe' nead are difficult to determine, it has been assumed that the fluid boundary conditions (fluid temperature) are similar to those in ti.e surge nozzle. In other words, the cold fluid entering the pressurizer has been assum'ed to fall to the bottom where it contacts the lower head.

, 3) The outside of the lower head, nozzle, and piping have been assumed to be fully insulated (no heat loss).

4) It has been assumed that the water between the thermal sleeve OD and nozzle ID is stagnant (no flow of water behind thermal sleeve).

5) The nozzle and pressurizer head are assumed to be at a uniform temperature

! at the beginning of each up ramp (peak) or down ramp (valley). This is l conservative since it maximizes the radial and - axial gradients -in the structure. Example: For an up ramp from '250F to 550F, the ramp begins l

with the nozzle and head at 250F.

6) The fluid temperature ramp rate (F/Hr) used in the analysis has been the maximum ramp rate at any time throughout the temperature change as defined in the design basis.

7) The stress indices from Table NB-3681(a)-1 for an "as welded transition" have been used in the evaluation of the safe end to elbow weld.

1 L

O 64

~ww ,-,e. -

g --- r ,-,,,,,n, ,, , , - ,,v, y ,,-,,.,-ww,w,,,,w,,-,,,,,,,nnw--exnw-,-y ,--e-, w s-,v,,r p- ,

The OBE seismic events have been assumed to occur at steady state O B) conditions and not at the point of maximum or minimum transient stress.

Even if an evea' were to occur during a time of maximum transient stress, the effect on fatigue usage for only one occurrence of OBE would be minimal, 6.1.5. Thermal Analvm An axisymmetric heat transf** analysis using the finite element code ANSYS has been performed to obtain the temperature distributions in the nozzle, safe end, welds, and pressurizer head. The thermal tt ansients evaluated are those specified in the design basis and discussed in Section 6.1.5.1. The resulting nodal temperatures from the thermal analysis have been used as input to the stress analysis.

The finite element model of tne pressurizer surge t c:zle is shown in figure 6 2.

The nozzle, safe end, welds, and pressurizer head are represented by isoparametric quadrilateral thermal elements STIF 55. The required inputs for this element are four nodal points and material properties: thermal conductivity, density, and specific heat, 6.1.5.1. Selection of Transients The operating transients for the surge line (and nozzles) have been identified in the design basis for the surge line and discussed in Section 4,5. A review of the transients revealed a significant number of temperature fluctuations during each transient. The temperature fluctuations involved approximately 70 different peaks or valleys per heatup and cooldown transient. The fluctuations include temperature changes from approximately 50 to 450 degrects F. An example of one of these transients, HULA 1, is town in Figure 6 3. As previously stated, an evaluation of each peak and valley is not practical, therefore only a few cases are considered. These cases are referred to as ' base cases" and have been selected to insure all peaks and valleys have been enveloped. A summary of the base case parameters and resulting stresses is given in Table 61, 6.1.5.2. Thermd_BoundaryConditions The thermal bounoary conditions consist of convective heat transfer at the inside surfaces of the pipe elbow, safe end, thermal sleeve, and pressurizer head.

6-5

~ , . -

i Because of the low velocities (generally less than 0.5 ft/sec) associated with O the in and out- surges, the heat transfer is by natural convection and is caused by the difference in temperature between the metal surface and the reactor coolant fluid. The film coefficient versus Delta T is input in tabular form to the ANSYS thermal runs. ANSYS uses the actual surface to fluid Delta T at each time step (iteration) to determine the appropriate film coefficient. A sample of the film coefficients for the various regions of the model is given below.

REGION LOCATION FILMCOEJ.

(BTV/Hr Ft F) 1 LOWER HEAD 288

. 2* BEHIND THERMAL SLEEVE COLLAR 145 3* BEHIND THERMAL SLEEVE, WATER REGION 19 4* BEHIND THERMAL SLEEVE, CONTACT ROLL 225 _

5 SAFE END AND PIPE. BEYOND SLEEVE 465 Effective film coefficient, takes into account thermal resistance of thermal sleeve, contact pressure, and water behind sleeve.

O The outside surfaces of the pipe elbow, safe end, nozzle, and pressurizer head

.have been assumed to be fully insulated, in addition, for symmetry, the ends of the pipe elbow and head have been assumed to be insulated.

A summary of the thermal boundary conditions is shown in Figure 6 4, 6.1.5.3. Results of Thermal Analysis The results of the thermal analysis are in the form of nodal temperatures. These nodal temperatures are read into the thermal. stress analysis and provide the model-with the axial and radial thermal' gradients- that produce the thermal stress. The times at which the maximum gradients occur are used for the thermal stress analysis since they will produce the maximum stresses. To determine when the maximum gradients occur the ANSYS postprocessor, POST 26, has been used.

POST 26 provides a time history-of the gradients at defined locations for the duration of the transient. For the surge nozzle evaluation fifteen pairs of nodes have been used'to examine the thermal response (Delta T) of the structure.

The locations of these 15 node pairs are shown in Figure 6 5. Since all the 6-6 n--,-~-m, n ..---,-,--,,-s,,.nn, , - ~ ~ , , , ,,.n.,,

transient peaks and valleys are defined by a fluid temperature ramp, the maximum O thermal gradients (and maximum stress) occur at the end of the applied ramp.

6.1.6. Stress Analysis An axisymmetric stress analysis using the finite element code ANSYS has been performed to obtain the stress distribution in the pipe, safe end, nozzle, and pressurizer head. The loadings for the analysis are the nodal temperatures from the thermal analysis (Section 5.1.5) and pressures from the design basis.

6.1.6.1. Descriotion of Finite Elenent Modgl The finite element model used for the thermal analysis has also been used for the stress analysis. The only difference between the two models is the element type designation. The nozzle, safe cnd, welds, and pressurizer head are represented by isoparametric quadrilateral stress elements, STIF 42. The model representing these components is the same as that shown in Figure 6-2. The required inputs for this element are four nedal points and material properties: density, coefficient of thermal expansion, Young's Modulus, and Poisson's ratio.

6.1.6.2. Structural Boundary Conditions The structural boundary conditions applied were required to simulate those portions of the structure tha' had not been modeled. The end of the model representing the pressurizer head has been restrained from motion in the l

meridional direction (UY displacement - 0.0). This restraint simulates the restraint of the adjacent head material. The nodes at~ the end of the pipe (elbow) have been coupled together in the axial direction for symmetry. A summary of these boundary conditions is shown in Figure 6 6.

6.1.6.3. Selection of Transient Tig i As stated in Section 6.1.5.3, the selection of transient times for use in the stress analysis is depende3t upon the thermal gradients through the structure.

The thermal gradients cause differential growth between adjacent material which results in thermal stresses. The times at which the maximum radial- and axial gradients (Delta-T) occur have been evaluated for stress. The results of the analysis shows that the radial gradient was the major contributor for maximum stress and almost always occurs at the end of the defined thermal ramp.

6-7

- - - - . - . - - - - - n.,-.. - - , - - - . . - , _ , _y- - . . . , ----,,,---e --tw----- -yy, ---

6.1.6.4. Finite Element Stress Results O The results from the ANSYS stress runs were not in a format which can be directly compared to ASME code allowables. In order to get stresses compatible with the ASME code requirements it was necessary to use the ANSYS postprocessor POST 11.

POST 11 performs stress linearization by converting the non-linear through wall stress distributions into the stress components required for an ASME code evaluation: membrane stress, bending stress, and peak stress. The pertinent information about the linearization method.; and detailed input is given in Section 6.31 of the ANSYS users manual. Eleven stress classification lines (SCL) have been selected to evaluate stresses in the various regions of the model. The line locations are shown in Figure 6 7. After being combined, the linearized stresses have been used to compare to the ASME code limit for the range of primary plus secondary stress intensities (3*Sm limit). An exau.rle of the linearized thermal and pressure stresses for the various " base cases" evaluated is shown in Table 6 1. The tabulated stresses are those for the safe end-to-elbow weld, SCL 10.

Also contained in Table 6-1 are the maximum stresses for the different " base cases". The maximum stresses represent the stresses at the surface of the component and are given in the ANSYS element stress printout as the element surfaco stress. When combined, the maximum stresses are used in the evaluation for fatigue usage.

6.1.6.5. ASME Code Calculations The linearized thermal, pressure, and external load stresses for each peak and valley were multiplied by the appropriate stress indices and combined to obtain the total linearized stress. The linearized stresses for all peaks and valleys have been tabulated and the difference between the maximum and minimum linearized stresses have been used for comparison to the ASME code limit of 3Sm. A sample of the linearized stresses for HVIAl at the safe end-to-elbow weld is given in Table 6-2. When the 35m limit has been exceeded, NB 3228.5 " Simplified Elastic-Plastic Analysis" has been used to justify the stress conditions.

The maximum thermal, pressure, and external load stresses for each peak and valley have been multiplied by the appropriate stress indices or stress .

concentration factor and combined to obtain the total maximum stress. The L 6-8 w

maximum stresses for all peaks and valleys have been tabulated for evaluation of V fatigue. A sample of the maximum stresses for HUIAl at the safe end to-elbou weld is given in Table 6-3.

The fatigue evaluation took into account the number of cycles for each peak and valley, the maximum stress ranges, the linearized stress range associated with the maximum stress range, and the resulting Ke factor for the maximum stress range when the linearized stress range exceed the 35m allowable.

The maximum linearized stress and fatigue usage factor for both the stainless steel and carbon steel portions of the surge nozzle are given in Section 6.1.7.

All strestes and fatigue usage factors meet the ASME code limits.

6.1.7. Summary of Results and Conclusion A summary of results for the prossurizer surge nozzle evaluation is given in the following table. Although the 3Sm limit is exceeded for'both the carbon steel and stainless steel, the requirements of the ASME code have been satisfied by performing a " Simplified Clastic-Plastic Analysis" as defined in NB-3228.5 of the code.

PRESSURIZER SURGE N0ZZLE CALCULATED FATIGUE USAGE FACTORS LOCATION Oconee Oconee Oconee TMl-1 CR-3 ANO-1 Unit 1 Unit 2 Unit 3 SAFE END TO ELBOW WELD .40 .41 .41 .33 .32 .32 (STAINLESS STEEL)

N0ZZLE TO HEAD CORNER .35 .35 .35 .32 .32 .32 (CARBON STEEL)

In conclusion, the pressurizer surge nozzle, safe end, and safe end to-elbow weld meet the requirements for Class I components of the ASME code, Section III,1986 Edition with no Addenda for the revised design basis transients discussed in Section 4.5.

(O y 6-9

p 6.2. Hot lea Surae Nozzle V The purpose of this section is to describe the evaluation of the hot leg surge nozzle. The stress analysis of the nozzle and nozzle-to surge line weld has been performed using the finite element method as implemented by the *ANSYS" computer ,

code, reference 17. The loads used for evaluation were the thermal and pressure loads identified in the design basis transients for the surge line stratification (see Section 4.5). The acceptance criteria for the evaluation were the requirements for Class I components of the ASME B&PV code, Section 111, 1986 edition with no Addeno , reference 18. The nozzle and nozzle to surge line weld were evaluated using the detailed requirements of NB 3200 as permitted by NB-3600.

6.c.l. Geometry An axisymmetrical representation of a small segment of the surge line, hot leg surge nozzle, and hot leg is shown in Figure 6-8 with an effective radius for the sphere (hot leg) equal to 3.2 times the hot leg pipe radius. The 3.2 to 1 equivalent spherical vessel is a modeling technique recommended by reference 20.

Using the 3.2 factor instead of the more common 2.0, assures that the maximum pressure stress at the critical location in the nozzle is adequately predicted by the axisymmetric model. This modeling technique is conservative for predicting the membrane stress but is accurate for predicting the maximum stress in the critical locations for use in a fatigue analysis. This piping junction consists of a carbon steel nozzle welded to the carbon steel hot leg. Both the nozzle and hot leg are clad with stainless steel to prevent reactor coolant fluid from contacting the carbon steel bue metal.

6.2.2. Descriotion of loadinos The loadings on the hot leg surge nozzle consist of thermal gradients, internal pressure, and external piping loads.

The thermal gradients are caused by the various fluid temperature swings (peaks and valleys) associated with the in- and out- surges of fluid between the pressurizer (hot) and the hot leg pipe (cold). The surge line fluid becomes stratified near the hot leg nozzle producir.g circumferential temperature gradients and thermal striping which are in addition to the axisymn'etric (radial

. and longitudinal) temperature gradients produced by the transient. Also, the 6 10 e

\

l

i temperature differential between the surge line fluid and the hot leg fluid contribute to these temperature gradients. The temperature swings for the surge nozzle and hot leg fluids are defined in the design basis for the surge line.

j The thermal gradients and stresses due to these temperature swings are determined using the ANSYS finite element code.

The RC system pressure is applied to the internal surfaces of the nozzle and hot leg pipe. The pressures are defined in the design basis and the resulting stresses are determined using the ANSYS finite element code.

There are significant external loads developed due to the heating and cooling of the surge line as well as the stratification in the surge line. The external loads, forces and moments, for cach peak and valley are given in the documentation of the surge line analysis. The stresses due to these moments and forces are calculated using the ANSYS finite element code.

6.2.3. Discussion of Analysis In a typical fluid temperature spike, the top fluid in the nozzle will have a larger temperature change than the bottom fluid. Thus, for the determination of the radial and longitudinal temperature gradients and the associated thermal q

D stress, it is conservative to use an axisymmetric analysis with the top fluid as the fluid boundary. An axisymmetric thermal and thermal stress analysis has been performed using the ANSYS finite element' computer code. The transient thermal analysis consists of imposing time dependent boundary conditions (bulk fluid l

temperatures and heat transfer coefficients) on the finite element model. Nodal temperatures from the thermal analysis are stored on magnetic tape for each iteration (time step) of the transient. The ANSYS postprocessor " POST 26" uses the nodal temperature: to calculate Delta-T's between various locations in the structure. Tables of the Delta T's versus time for each transient are used to determine when the maximum and minimum stresses are likely to occur. The nodal i temperatures for each critical time step are input to the ANSYS stress routine-for the determination of stressas. The ANSYS postprocessor, ' POST 11", is used l

l to linearize the stresses at critical sections of the. structure. Stresses due l to pressure and resultant external force (along the- nozzle axis) are also determined at t.he critical sections using ANSYS and POSTil.

i 6 11

q Due to the large number of temperature swings (peaks and valleys) associated with C/ the thermal stratification transients it was not practical to evaluate each peak and valley as an individual case. Instead, a few ' base cases' were created to envelop the large number of identified peaks and valleys. The base cases were chosen using the following parameters; the maximum instantaneous ramp rate for the top fluid temperature excursion in the stratified nozzle, the hot leg temperature, the convective heat transfer coefficients for the nozz1t and hot leg, and the top fluid temperature change (Delta T) between a peak and valley in the nozzle. The temperature distribution and resulting thermal stresses for each of the base cases were determined using the procedure described above. A summary of the parameters used in the base cases is given in Table 6 4. Parameters describing the actual transients identified in the design basis were used to determine a representative base case for each peak and valley which will now approximate the actual transient. The linearized and maximum thermal stresses from the chosen " base case

• were used directly for combining with stresses due to pressure, resultant external force, and non axisymmetic load stresses.

Pressure stresses for a base case with an internal pressure of 1000 psi were e determined using ANSYS, The pressure stresses for each peak and valley were k determined by multiplying the stresses from the base case by the ratio of the actual pressure for the peak or valley from the design basis to the pressure used in the pressure base case.

5 The stresses due to a axial load of 10 lbs were determined using ANSYS. The axial locd stresses for each peak and valley were determined by multiplying the stresses from the base case by the ratio of the axial force from the surge line evaluation at the time nearest the peak or valley to the force used in the base Case.

As described above, the thermal analysis performed is axisymmetric in that it assumes that the top fluid completely fills the nozzle and creates the two-dimensional axisymmetrical temperature fields in the nozzle for the various thermal transients. The task now is to determine the additional stresses due to circumferential temperature gradients produced by the fluid stratification in the nozzle. The stresses due to this fluid stratification will conservatively be assumed to occur at the time of maximum thermal stresses due to the radial and longitudinal temperature gradients.

--__-_____-___-____-_-_____J

i i

i The stresses due to thermal stratification are determined for two base cases by use of the ANSYS harmonic element $TIF 25. This element is used for two-l dimensional modeling of an axisymmetric structure with nonaxisymmetric loading, In the case being considered, the nonaxisymmetric loading is the temperature f

! field in the nozzle which varies in the circumferential direction as well as in the radial and axial directions. The stresses due to a circumferential temperature gradient are independent of the radial and axial gradients. These l stresses are primarily a function of the temperature difference between the top and bottom fluid and the transition zone between the two fluid temperatures.

The circumferential temperature gradient is approximated by assuming the top and

! bottom of the nozzle are at a steady state condition for a thermal peak and valley, respectively. The transition between these two temperature fields is assumed to be linear over the same l" height of nozzle that contains the fluid interface zone between the hot and cold fluid. From the design basis, the centerline elevation of this-interface zone in the hot leg nozzle, relative to the centerline of the nozzle, varies / rom 0.0" to -3.88" during the various PV temperature excursions.- The elevation (Z) of the interface zone is assumed to I be at 0.0" for actual elevations of 2.0" < -Z r 0.0" and --3.88" for actual

(

elevations of 3.88" s Z s -2.0". Thus, only two thermal stratification load -

base cases are required.

The two stratification base cases used a 229F temperature differential between the hot and cold fluid. Therefore, the stress due to the base case circumferential temperature gradient can be determined by subtracting the steady state stress due to the radial and axial temperature gradients from the combined stress due to radial, axial, and circumferential temperature gradients (from harmonic element results which included- the same steady state temperstures as were used in the axisymmetric load).

The thermal stratification stresses due to the circumferential temperature gradient for each peak and valley were determined by multiplying the stresses from the appropriate circumferential temperature gradient base case by the ratio of the actual stratification Delta T for the peak or valley from the design basis to the Delta T used in the stratification base case (229F).

6-13

-,,--w-.-e.,--,nn-r--2,v.c m ,-,-,w,m_m.m.~ , . - . . . , w. ~ ~ . . ,1w-,,-w.,.., --.-,,-w-, ,.r.m , ~ . , .-y -v- -- +..c,- + - , . , . . , , ,t

1 Stresses for a base case nozzle bending moment of 10' in lbs (in the plane of the O surge line) were also determined by using the ANSYS harmonic element, STIF 25.

The bending moment stresses for each peak and valley were determined by multiplying the stresses from the base case by the ratio of the moment from the surge line evaluation at the time nearest the peak or valley to the moment used in the base case.

The thermal, stratification, pressure, and external load stresses were multiplied by the appropriate stress indices (Table NB 3681(a) 1) or stress concentration factors and then combined for determination of maximum stress and fatigue usage.

The results are given in Section 6.2.9.

6.2.4. List of Assumotions/Inouts Used in Analysis

1) The surge line nozzle to hot leg junction is a 3D structure and the thermal stratification in the nozzle produces non axisymmetric loads. Two i significant assumptions are necessary in order to analyze this nozzle using a 20 finite element model, a) The 30 structure can be approximated as a nozzle attached to a sphere whose radius is 3.2 times the radius of the hot leg. This assures that the pressure stress in the model will be equivalent to the maximum pressure stress in the actual structure at the critical

. location. The thermal stress from the model due to temperature gradients is approximately equal to those in the actual structure since thermal stress is not a strong function of the radius of the sphere, b) The circumferential temperature gradient is approximated by assuming the top and bottom of the nozzle is at a steady state condition for a thermal peak and valley, respectively. The transition between these two temperature fields is assumed to occur over the same 1" height of nozzle that contains the fluid interface zone between the hot and cold fluid. This is a conservative assumption since heat conduction in the circumferential direction of the nozzle will increase the height of the transition zone in the metal which would tend to reduce the thermal stresses.

O 6-14

l c 2) The outside of surge line nozzle and hot leg are assumed to be fully

( insulated (no heet loss).

3) The surge line nozzle and hot leg are assumed to be at a steady state condition at the beginning of each up ramp (peak) or down ramp (valley).

This is a conservative assumption since it maximizes the radial and axial gradients in the structure for each peak or valley. >

4) The fluid temperature ramp rate (F/Hr) used in the analysis in determining the applicable base case is the maximum ramp rate at any time throughout thetemperaturechange(PV)asdefinedinthedesignbasis.

5) The transition between the nozzle and het leg fluid' temperatures is assumed to be a step change occurring at the intersection of the nozzle and hot leg pipe. This is a conservative assumption as a more gradual transition will actually occur which would reduce the thermal stresses in this region of the nozzle.

6) The stress indices from Table NB-3681(a)-1 for an "as welded transition" are used in the evaluation of nozzle-to surge line weld. This conservatively assumes the field weld was not ground flush on the inside b surface of the nozzle.

7) The OBE seismic events are assumed to occur at steady state conditions and not at the point of maximum or minimum transient stress. Even if an event were to occur during a time of maximum transient stress, the effect on '

fatigue usage for only one occurrence of OBE would be minimal.

6.2.5. Thermal Analysis of Axisymmetric loads A3 axisymmetric heat transfer analysis using the finite element code ANSYS is performed to obtain the temperature distributions in the surge line nozzle and hot leg. The thermal transients evaluated are those specified in the design basis and discussed in Section 6.2.5.1. The resulting nodal temperatures from the thermal analysis will be used as input to the stress analysis.

The finite element model of a small segment of the surge line, surge nozzle, and hot leg is shown in Figure 6 9.- These components are represented by isoparametric quadrilateral thermal elements, STIF 55. The required inputs for ,

6-15

this element are four nodal points and material properties: thermal conductivity.

l density, and specific heat.

6.2.5.1. Selection of Transients The operating transients for the surge line (and nozzles) are identified in the design basis for the surge line and discussed in Section 4.5. A review of the transients revealed a significant number of temperature fluctuations during each transient. The temperature fluctuations involved approximately 40 different peaks or valleys per heatup and cooldown transient. The fluctuations include temperature changes (Delta T) with magnitudes ranging from approximately 50 to 400 F. An example of one of these transients, HU1A1, is shown in Figure 6 10.

As previously stated, an evaluation of each peak and valley is not practical, therefore only a few cases are considered. These cases are referred to as " base cases" and are selected to insure all peaks and valleys are enveloped. A summary of the base case parameters and resulting stresses is given in Table 6 4, 6.2.5.2. Thermal Boundary Conditions The thermal boundary conditions consist of convective heat transfer at the inside r surfaces of the model. Depending on the flow velocity in the nozzle and hot leg, either free or forced convection may be the predominant mode of heat transfer l between the fluid and metal surfaces. For natural convection, the heat transfer l

1s caused, primarily, by the difference in temperature between the metal surface l

and the reactor coolant fluid. The film coefficient versus Delta-T is input-in tabular form to the ANSYS thermal runs. ANSYS uses the actual surface to fluid l Delta T at each time step (iteration) to determine the appropriate film coefficient. For forced convection, the film coefficient is constant for a given fluid velocity, temperature, and geometry. The film coefficient used in the analysis is the maximum of the coefficients for free or forced convection. A sample of the film coefficients for the two regions of the model is given below.

REGION' FILM COEFFICIENT-(BTU /HR FT2 .p)

FREE FORCED-N0ZZLE 140 185 HOT LEG 115 3000 6-16

The outside surfaces of the model are assumed to be fully insulated. In  :

O addition, for symmetry, the ends of the model are assumed to be adiabatic surfaces. .

j l The transition between the nozzle and hot leg fluid temperatures is conservatively assumed to be a step change at the intersection of the nozzle and i hot leg pipe, i 6.2.5.3. Results of Thermal Analysis -

l The results of the thermal analysis are in the form of nodal temperatures. These nodal temperatures are read into the thermal stress analysis and provide the model with the axial and radial thermal' gradients that' produce the thermal' stress. The times at which the maximum gradients occur are used for the thermal stress analysis since they are likely to produce the, maximum-stresses. To determine when the maximum gradients occur the ANSYS postprocessor, POST 26, is used. POST 26 provides a time history of the gradients at defined locations for the duration of the transient. For the surge nozzle evaluation fifteen pairs of f nodes were used to examine the thermal response (Delta T) of the structure.- The locations of these'15 node pairs are'shown in figure 6 11. ,

O figure 6 12 shows the temperature contours at an extreme Delta T time point of a typical fluid temperature spike. <

4

{ 1 6. Stress Analysis of Axisymmetric Loads ,

An axisymmetric stress analysis using the finite- element code ANSYS has been performed to obtain the stress distribution in the model' for the base case axisymmetric loadings. The loadings for the analysis are the nodal temperatures from the thermal analysis (Section 6.2.5), a unit pressure load-(1000 psi), and 5

a unit axial force (10 lbs). .

6.2.6.1. Descriotion of Finite Element Model The finite element model used for the thermal' analysis was'_also. used for the stress analysis. The only difference between the two models is the element type designation. _The ST1F. 55 thermal element is replaced with an isoparametric h quadrilateral stress element, STIF 42. The required inputs for this element are four nodal points . and material properties: coefficient of thermal expansion,

- modulus of elasticity, and Poisson's ratio.

- - 6 17

. a r

,-wr, .-in.- .n..- ,.n a +~,w . . ~,._-n-, na.n,.r,6-- w., -,,.w. a

l 6.2.6.2. Structural Boundary Conditions i O The structural boundary conditions applied are required to simulate those portions of the structure that are not modeled. The end of the model ,

representing the hot leg was restrained from motion in the meridional direction (UY displacement 0.0). This restraint simulates the restraint of the adjacent hot leg pipe material. The end of the surge line segment is assumed to be free.

The location of this free boundary condition is sufficiently remote from the nozzle to surge line weld such that any stress ?nduced by the assumed boundary condition will have attenuated to a negligible value at this critical section.

6.2.6.3. Selection of Transient Times As stated in Section 6.2.5.3, the selection of transient times for use in the stress analysis is dependent upon the thermal gradients through the structure, The thermal gradients cause differential growth between adjacent material which results in thermal stresses. The times at which the maximum radial and axial gradients (Delta-T) occur are evaluated for stress.

6.2.6.4. Finitt, Element Stress Results The results*from the ANSYS stress runs are not in a format which can be directly l

compared to ASME code allowables, in order to get stresses compatible with the ASME code requirements it is necessary to use the ANSYS postprocessor POST 11.

l POST 11 performs stress linearization by converting the non linear through wall stress distributions into the stress components required for an ASME code evaluation: membrane stress, bending stress, and peak stress. The pertinent information about the linearization methods and detailed input is given in Section 6.31 of the ANSYS users manual. Eleven stress classification lines (SCL) were selected to evaluate stresses in the various regions of the model. The line locations are shown in Figure 6 13. The sum of the linearized stresses for a given load set is used to compare to the ASME code limit for the range of primary plus secondary stress intensities (35m limit). An example of the linearized and maximum thermal stresses for the various ' base cases" evaluated is shown in Table 6 4. The tabulated stresses are those for the nozzle to surge line weld.

Also contained 1.'. Table 6-4 are the maximum stresses for the different ' base cases'. The maximum stresses represent the stresses at the surface of the 6 18

l component and are given in the ANSYS element stress printout as the element -

l surface stress. The sum of the maximum stresses for a given load set is used in 3

the evaluation for fatigue usage.

6.2.7. Stress Analysis of Non agjsymetric loads

! A non axisymetric stress analysis using the finite element code ANSYS has been performed to obtain the stress distribution in the m' del for the base. case non-axisymetric loadings. The loadings for the analysis are the circumferential nodal temperature gradients for two stratification cases and a _ nozzle bending moment as described in Section 6.2.3.

6.2.7.1. Descriotion of Finite Element Model The finite element stress model used for the axisymetric loads was also used for the stress model for the non axisymmetric loads. The only difference between the two models is the element type designation.- The STIF 42 element is replaced with a harmonic element, STif 25. The required inputs for this element are four nodal points and constant material properties: coefficient of - thermal expansion, modulus of elasticity, and Poisson's ratio.

6.2.7.2. Structural Boundary conditions The structural boundary conditions are the same as was used for the axisymetric loads in Section 6.2.6.2.

6.2.7.3. ANSYS Load Sten Dat.A 1

A harmonic element model requires the load to be input as a series of harmonic

' functions (Fourier series). The ANSYS preprocessor PREP 6 is used to generate the -

I fourier series for the stratification temperature fields described in Section 6.2.3. Stresses were obtained for all the odd numbered modes up through mode number 33. These stress modes were then combined using the ANSYS postprocessor POST 29. The unit (10 6 in-lb)nozzlebendingmomentwasappliedasdescribedin-Section 2.2.5, Case C of the ANSYS user's manual. This bending _ moment is -

represented by applying peak axial (nozzle) force values at the end of _ the ,

nozzle. The load varies as a first harmonic wave (MODE - 1) with a cosine symmetry condition (ISYM = 1).

' 19

= , - , - - ,,,+-,2,--,,~.,-,%m vw w # w -,..-m.m ,: , . ~

n 6.2.7.4. Finite Element Stress Results I

! The stresses output from POST 29 were linearized at critical sections of the model. The results were then combined with linearized stresses from other loads and compared to the ASME code allowables as described in Section 6.2.8.

6.2.8. ASME Code Calculations The linearized thermal, stratification, pressure, and external load stresses for each peak and valley were combined to obtain the total linearized stress. The linearized stresses for all peaks and valleys were tabulated and the difference between the maximum and minimum linearized stresses was used for comparison to the ASME code limit of 35m. When the 3Sm limit was exceeded, NB 3228.5

' Simplified Elastic Plastic Analysis" was used to justify the stress conditions.

The maximum thermal, stratification, pressure, and external load stresses for each peak and valley were multiplied by the appropriate stress indices or stress concentration factor and combined to obtain the total maximum stress. The

maximum stresses for all peaks and valleys were. tabulated for evaluation- of fatigue. A sample of the linearized and maximum stresses at the nozzle to surge -

line weld with the associated fatigue usage for a typical PV is given in Table 6 5. The fatigue evaluation took into account the number of cycles for each peak and valley, the maximum stress ranges, the linearized stress range associated l with the maximum stress range, and the resulting Ke factor for the maximum stress -

range when the linearized stress range exceed the 35m allowable.

l Fatigue usage due to thermal striping on the stainless steel regions of the nozzle is conservatively assumed to be equal to that calculated for the surge

- line.

The maximum linearized stress and fatigue usage factor for both the stainless steel and carbon steel portions of the surge nozzle are given in Section 6.2.9.

All requirements of the ASME code are met.

6.2.9. Summary of Results and Conclusion A summary of results for the hot leg surge nozzle evaluation is given in the following table. Although the 35m limit is exceeded for both the carbon steel and stainless steel, the requi. nents of the ASME code were satisfied by O S-?0 l

l lW

-..,.m. ,, .-,.,-g , -.- ,,...mn,, , . , , . , , .,w,,,..,,, , , , . - ., -.._-nr,n,_w,n

, , , , .n..a -,e.,,n e,na ,. w,y..,-,,-,.. 4mm . - - , . - - . , . , , , ,

l performing a " Simplified Elastic Plastic Analysis' as defined in NB 3228.5 of the code.

HOT LEG SURGE N0ZZLE CALCULATED FATIGUE USAGE FACTORS LOCATION Oconee Oconee Oconee THI-l CR-3 ANO 1 Unit 1 Unit 2 Unit 3 N0ZZLE TO SVRGE LINE WELD .26 .26 .25 .19 .19 .20 (STAINLESS STEEL)

N0ZZLE TO HOT LEG CORNER .59 .59 .59 .41 .42 .42 l (CARBON STEEL) l END OF N0ZZLE TAPER .42 .42 .42 .62 .62 .62 (CARBON STEEL) _

in conclusion, the hot leg surge nozzle and nozzle to surge line weld meet the requirements for Class I components of the ASME code, Section Ill,1986 Edition with no Addenda for the revised design basis transients discussed in Section 4.5.

O 6-21 9

--,-,+e . . . . , . , . -- , ,-,, , -- - , -

l '

TABLE 6-1 BASE CASES STRESS LINE 10, INSIDE SURFACE, SAFE END-TO-ELBOW WELD TRANSIENT INFORMATION STRESSES Base Temperatures Ramp Press Linearized Maximum (el. 376)

Run, (F) Rate S12 S23 S31 S12 S23 S31 BR # Strt End Delta F/Hr (psi) (ksi) (ksi) (ksi) (ksi) (ksi) (ksi)

============== ====== ===== ==============

RUla 80 482 402 900 11.1 ~6.2 -4.9 14.7 ~7.7 ~7.0 RUlb 80 446 366 900 10.9 -6.0 -4.9 14.5 -7.4 -7.1 RUlc 80 373 293 900 10.2 -5.3 -4.9 13.7 -6.5 -7.2 RUld 80 309 229 900 9.5 -4.6 -4.9 12.8 -5.7 -7.1 Rule 80 253 173 900 8.7 -3.0 -4.9 11.9 ~4.9 ~7.0 RUlf 80 209 129 900 7.7 -3.0 -4.7 10.7 -4.0 -6.7 RU19 80 172 92 900 6.5 -2.3 -4.2 9.2 -3.1 -6.1 RU2a 150 500 350 1500 15.8 -7.6 -8.1 21 4 -9.6 -11.6 RU2b 150 443 293 1500 15.1 ~6.9 -8.2 20.4 ~8.7 -11.7 RU2c 150 383 233 1500 13.8 -5.8 -8.0 18.9 -7.5 -11.4 RU2d 150 327 177 1500 12.2 -4.6 -7.6 16.9 ~6.1 -10.8 RU2e 150 285 135 1500 10.7 -3.7 -7.0 15.0 -4.9 -10.1 RU2f 150 248 98 1500 8.7 -2.7 -6.0 12.4 -3.7 -8.7 RU2g 150 220 70 1500 6.7 -1.9 ~4.9 9.8 -2.6 -7.2

(~) RU3a 250 550 300 1000 11.1 -5.8 -5.3 14.3 -7.2 -7.6

\l RU3b 250 484 234 1000 10.5 -5.1 ~5.4 14.1 -6.4 ~7.7 RU3c 250 430 180 1000 9.7 -4.2 -5.4 13.2 ~5.5 -7.7 RU3d 250 382 132 1000 8.5 -3.3 -5.2 11.7 -4.3 -7.4 RU3e 250 346 96 1000 7.2 -2.5 -4.7 10.1 -3.3 -6.8 RU3f 250 316 66 1000 5.6 -1.7 -3.9 8.0 -2.3 -5.7 RU3g 250 292 42 1000 3.8 -1.0 -2.8 5.6 -1.4 -4.2 RU4a 250 482 232 2000 16.4 -6.3 -10.1 22.6 -8.2 -14.4 RU4b 250 431 181 2000 14.6 -5.1 -9.5 20.3 -6.8 -13.5 RU4c 250 389 139 2000 12.4 -3.9 -8.5 17.5 ~5.3 -12.2 RU4d 250 352 102 2000 9.8 -2.8 -7.0 14.1 ~3.9 -10.2 HU4e 250 320 70 2000 7.0 -1.8 -5.2 10.3 -2.6 -7.7 RU5a 400 650 250 2000 16.8 -6.8 -10.0 22.9 -8.8 -14.1 RU5b 400 635 235 2000 16.5 -6.5 -10.0 22.5 -8.4 -14.1 RU5c 400 580 180 2000 14.6 ~5.2 -9.4 20.3 -7.0 -13.3 RU5d 400 535 135 2000 12.6 -4.0 -8.6 17.6 -5.4 -12.2 RU5e 400 500 100 2000 10.2 -2.9 -7.3 14.6- -4.1 -10.5 RU6a 500 650 150 1000 8.7 -3.7 -5.0 11.9 -4.8 -7.1

, RU6b 500 611 111 1000 7.8 -3.0 -4.8 10.7 -3.8 -6.9 RU6c 500 579 79 1000 6.5 -2.1 -4.3 9.1 -2.9 -6.2 RU6d 500 555 55 1000 5.0 -1.4 -3.6 7.1 -1.9 -5.2 RU7a 500 650 150 1500 11.2 -4.2 -7.0 15.4 -5.4 -10.0 RU7b 500 624 124 1500 10.2 -3.5 -6.7 14.2 -4.7 -9.5 RU7c 500 590 90 1500 8.4 -2.5 -5.8 11.8 -3.4 -8.4 em S550l550 550 0 l NONE l l -0.3 0.2 0.1 l -0.4 0.3 0.1 l m

6-22

_ _ _ _ _ _ J

t t

TABLE 6-1 BASE CASES STRESS ~?NE 10, INSIDE SURFACE, SAFE END-TO-ELBOW WELD TRANSIENT INFORMATION STRESSES Base Temperatures Ramp Press Linnarized Maximum (el. 376)

Run, (F) Rate S12 .S23 S31 812 S23 S31 BR # Strt End Delta F/Hr (psi) (ksi) (ksi) (ksi)

(ksi) (ksi) (ksi).

S450 450 450 0 NONE -0.3 0.2 0.1 -0.3 0.2 0.1 S3C0 300 300 0 NONE -0.2 0.1 0.1 -0.2 0.1 0.1 SS70 70 70 0 NONE 0.0 0.0 0.0 0.0 0.0 0.0 RDia 550 300 ~250 -800 -9.7 5.2 4.5 -13.0 6.5 6.5 RDib 550 332 -218 -800 -9.4 4.8 4.6 -12.6 6.1 6.5 RDic 550 386 -164 -800 -8.6 4.1 4.5 -11.5 5.1 6.4 RDid 550 432 -118 ~800 -7.6 3.2- 4.3 -10.3 4.2 6.1 RDie 550 468 -82 -800 -6.4 2.4 3.9' -8.8 3.2 5.6 RDif 550 493 -57 -800 -5.1 1.8 3.4 -7.2 2.4 4 ~. 8 RDig 550 514 -36 -800 -3.6 1.1 2.5 -5.2 1.6 3.6 RD2a 482 160 -322 -4000 -28.1 9.8 18.3 -39.7 13.1 26.6 RD2b 482 218 -264 -4000 -25.0 8.1 16.9. -35.5 11.1 24.4 '

RD2c 482 276 -206 -4000 -21.1 6.4 14.7 -30.2 8.9 21.3 RD2d 482- 321 -161 -4000 -17.3 4.9 12.4 -25.0 6.9 18.1

{, RD3a 450 100 ~350 -3000 -25.2 9.8 15.4 -35.2 12.9 22.3 RD3b 450 135 -315 -3000 -24.0- 9.1 15.0 -33.7 12.0 21.7 RD3e 450 198 -252 -3000' -21.4 7.5 13.9 ~30.2 10.1 20.1 RD3d 450 254 -196 -3000 -18.4 6.0 12.5 -26.2 8.2 18.0 RD3e 450 296 -154 -3000 -15.6 4.7 10.9 -22.3 6.5 15.8 RD3f 450 338 -112 -3000 -11.9 3.3 8.6 -17.3 4.7 12.6 RD3g 450 373_ 3000 -8.2 2.1 6.1 -12.2 3.1- 9.1 RD4a 450 100 -350 -1700 -17.7 8.3 9.4 -24.2 10.5 13.7 RD4b 450 _145 -305 -1700 -17.0 7.6 9.4 -23.3 9.8 13.5 RD4c 450 215 -235 -1700 -15.3 6.3 9.0 -21 '.' 2 8.3 12.9 RD4d 450 270 -180 -1700 -13.6 5.2 8.5 -18.9 6.8 12.1 I RD4e 450 315 -135 -1700 -11.6 4. 0- 7.6 -16.3 5.4 10.9

! RD4f 450 350 -100 -1700 -9.5 3.0 6.5 -13.4 4.1 9.3 RD4g 450 380 -1700 -7.1 2.1 5.0 -10.2 2.9 7.3 RD4h 450 405 1700 -4.6 1.2- 3.3 -6.8 1.8 5.0

! RD5a 300 100 -200 -700 -7.8- 3.9 3.9 -10.61 4.9 5.7 RD5b 300 137 -163 -700 ~7.4 3.4 3.9- -10.1 4.4 5.7 .

RD5c 300 183 -117 -700 -6.5 2.7 3.8 -9.0 3.6 5.4 l RD5d 300 217 -83 -700 -5.5 2.1 3.4 -7.7 2.8 4.9 RD5e 300 243 -57 -700 -4.4 1.5 2.9 -6.2 2.1 4.1 PRS l l l2200 l -9.1- -3.1l 12.2 l-10.3 -2.2 12.5 l h:#

u o

I 6-23

• 1 O O

~

O TABLE 6-2: L!WEAaIZED staESSES, TRAustEst 1A1 Stat $$ LlaE 10, tuslDE SLMACE, SME EaD-TO-ELBCN WELD PV W ' Temperatures Mea Thermal Stresses Pressure Stresses Pipirg Stresses Lineerited Stress s emp lSese pW Mammt Stress (see sotes) strt Erd Detta mete stan 512 523 531 Press 512 S23 S31 load Mrtryr) Mc/t 512 523 531 (F) (F) (F) (T/ar) Used (tsi) (kst) (ksi) (psi) (tsi) (ksi) (ksi) pv i (Iri-lbs) (tsi) (tsi) (ksi) (tsi)

Pv1 70 70 0 0 SS70 0.0 0.0 0.0 0 0.0 0.0 0.0 1 24024 0.3 0.4 0.4 0.0 Pv2 70 482 412 848 aule 11.1 -6.2 -4.9 578 -2.4 -0.8 3.2 2 2457310 29.8 46.8 31.1 -1.7 Pv3 482 163 -319 -3954 WD2e -28.1 9.8 18.3 578 -I.4 -0.8 3.2 3 522099 6.3 -22.4 17.1 21.5 Pv4 163 482 319 758 au3e 11.1 -5.8 -5.3 578 -2.4 -0.8 3.2 4 2508422 30.4 47.6 32.3 -2.1 PV5 482 97 -385 -3849 aD2e -28.1 9.8 18.3 578 -2.4 -0.8 3.2 5 280905 3.4 -26.1 13.3 21.5

• v6 97 240 143 692 au1f 7.7 -3.0 -4.7 578 -2.4 -0.8 3.2 6 2%56005 30.1 43.9 34.7 -1.5 Pv7 240 104 -1M -292 RD5e -6.5 2.7 3.8 578 -2.4 -0.8 3.2 6 2486005 30.1 29.7 40.4 T.0 Pv8 104 338 234 1305 au2c 13.8 -5.8 -8.0 578 -2.4 -0.8 3.2 6 2486005 30.1 50.0 31.9 -4.8 PW9 338 101 -237 -2370 aD3e -21.4 T.5 13.9 578 -2.4 -0.8 3.2 7 41122T 5.0 -17.4 13.1 17.1 PV10 101 269 168 1218 au2d 12.2 -4.6 -7.6 578 -2.4 -0.8 3.2 8 1829208 22.2 38.2 22.9 -4.4 Pv11 269 135 -134 -651 eD5e -6.5 2.7 3.8 578 -2.4 -0.8 3.2 9 483188 5.9 -1.4 9.4 7.0 Pv12 135 178 43 1538 auze 7 -1.9 -4.9 578 -2.4 -0.8 3.2 9 483188 5.9 11.8 4.8 -1.7 Pv13 178 110 -68 -724 eD5d %5 2.1 3.4 578 -2.4 -0.8 3.2 11 28077T 3.4 -3.5 5.6 4.6 Pv14 110 248 138 1390 au2e 10.7 -3.7 -7.0 578 -2.4 -0.8 3.2 12 2195571 26.6 42.4 29.5 -3.8 Pv15 248 114 -134 -376 eD5c -6.5 2.7 3.8 578 -2.4 -0.8 3.2 ?3 429664 5.2 -2.2 8.5 7.0 Pvt 6 114 266 152 1717 au2d 12.2 -4.6 - T.6 578 -2.4 -0.8 3.2 14 1382962 16.8 31.3 16.0 -4.4 Pv17 266 117 -149 -666 aD5b -7.4 3.4 3.9 578 -2.4 -0.8 3.2 15 378692 4.6 -3.9 8.5 7.1 Pv18 11T 255 138 8 73 autf 7.7 -3.0 -4.7 578 -2.4 -0.8 3.2 16 1648571 20.0 30.9 21.7 -1.5 Pv19 255 130 -125 -428 Rc5c -6.5 2.7 3.8 578 -2.4 -0.8 3.2 17 440821 5.3 -2.1 8.7 T.0 Pv20 130 321 191 1192 au2d 12.2 -4.6 -T.6 578 -2.4 -0.8 3.2 18 1695162 20.5 36.1 20.9 -4.4 cs Pv21 321 140 -181 -823 RD5e -7.8 3.9 3.9 578 -2.4 -0.8 3.2 19 3957M 4.8 -4.1 9.2 T.1 au2s 15.8 -7.6 -8.1 578 -2.4 -0.8 3.2 23 39825 1.7 15.6 -6.2 -4.9 h

=

Pv22 Pv23 140 478 338 478 305 -173 1294

-352 acic -8.6 4.1 4.5 578 -2.4 -0.8 3.2 24 1536736 18.6 12.8 27.1 7.7 Pv24 305 380 75 1T67 aute 7.0 -1.8 -5.2 578 -2.4 -0.8 3.2 24 15M7M ts.6 28.4 21.2 -2.0 Pv25 380 253 -127 -395 eD5e -6.5 2.7 3.8 578 -2.4 -0.8 3.2 24 15MT36 18.6 14.9 25.T T.0 Pv26 253 300 47 1572 aute 7.0 -1.8 -3.2 578 -2.4 -0.8 3.2 24 15M736 18.6 28.4 21.2 -2.0 Pv27 300 231 -69 -139 aD5e -4.4 1.5 2.9 578 -2.4 -0.8 3.2 25 306030 3.7 -2.0 5.4 6.1 Pv28 231 431 200 873 au3c 9.7 -4.2 -5.4 578 -2.4 -0.8 3.2 L 13G2842 15.8 27.5 15.2 -2.2 Pv29 431 243 -188 -807 acte -8.6 4.1 4.5 578 -2.4 -0.8 3.2 27 406195 4.9 -4.T 9.6 7.7 Pv30 243 328 85 1466 autd 9.8 -2.8 -7.0 578 -2.4 -0.8 3.2 28 9363T4 11.3 21.9 10.9 -3.8 Pv31 328 234 -94 -197 eD5d -5.5 2.1 3.4 578 -2.4 -0.8 3.2 29 133506 1.6 -3.8 3.4 6.6 Pv32 234 4TT 243 884 au3b 10.5 -5.1 -5.4 578 -2.4 -0.8 3.2 31 428891 5.2 14.8 0.7 -2.2 Pv33 4TT 234 -243 -350 ante -9.7 5.2 4.5 578 -2.4 -0.8 3.2 30 1689142 20.5 14.1 30.6 7.7 Pv34 234 418 184 1357 autb 14.6 -5.1 -9.5 578 -2.4 -0.8 3.2 32 1354047 16.4 33.2 15.1 -6.3 PY35 418 344 - 74 -1400 mD49 -T.1 2.1 3.0 578 -2.4 -0.8 3.2 33 929878 11.3 4.9 15.T 8.2 Pv36 344 429 85 1241 au3e 7.2 -2.5 -4.7 578 -2.4 -0.8 3.2 34 1248528 ?5.1 24.2 16.0 -1.5 Pv37 429 276 -153 -977 eD1c -8.6 4.1 4.5 659 -2.7 -0.9 3.7 35 458860 5.6 -4.2 10.3 8.2 Pv38 276 411 135 1751 aute 12.4 -3.9 -8.5 708 -2.9 -1.0 3.9 36 1585130 19.2 34.1 19.7 -4.6 Pv39 411 282 -129 -848 eDid -T.6 3.2 4.3 939 -3.9 -1.3 5.2 37 194453 2.4 -8.5 4.9 9.5 Pv40 282 591 309 1058 aula 11.1 -5.8 -5.3 1543 -6.4 -2.2 8.6 38 2092208 25.3 37.2 24.5 3.3 Pv41 591 524 -67 -2433 aD39 -8.2 2.1 6.1 1551 -6.4 -2.2 8.6 39 1335898 16.2 20.6 14.7 Pv42 524 584 60 659 au6d 5.0 -1.4 -3.6 1587 -6.6 -2.2 8.P 40 1799577 2ti 24.3 5.2 Pv43 584 469 -115 -494 aD1d -7.6 3.2 4.3 1655 -6.8 -2.3 9 41 1446453 1 ' 3 13.5 Pv44 469 546 77 1406 ru7e 8.4 -2.5 -5.8 1674 -6.9 -2.4 9.3 42 **72581 R

• 3.5 Pv45 546 376 -170 -946 eDic -8.6 4.1 4.5 1765 -7.3 -2.5 9.8 43 - GS1 9 4.3

O r~ .

J 'N ,

l TA8tE 6-2: LINEARIZED STRESSFS, tamrSIE1ft 1A1 STRESS tlWE 10, INSIDE SURF 8tE, SAFE END-TO-EL9OW WELD themet Stresses [ Pressure Stresses [ Piping Stresses tifwerited Stress PV W temperatures Max (see motes)

Reap Sese  ! pipe pseent Stress Rm 512 323 S3' Press 512 SZ3 531 tend pr(ryr) mc/I S12 S23 $31 Strt End Dette sete pe # (ksi) (ksi) (ksi)

(F) (F) (F) (F/pr) 47sel (ksi) (ksi) (ksi) (psi) (ksi) (ksi) (ksi) (Ift-its) (ksi) 10.2 -2.9 -T.3 1778 - T.4 -2.5 9.9 44 7567425 19.0 27.1 15.9 2.6 Pv46 376 468 92 2209 RU5e 4.3 -10.2 6.0 15.2 Pv47 468 344 -124 -558 ecto -T.6 3.2 4.3 1968 -8.1 -2.8 10.9 45 356343 12.4 -8.5 2020 -8.4 -2.8 11.2 46 1805393 21.9 32.0 21.2 2.7 Pv43 344 481 137 1670 RU4e -3.9 2062 -8.5 11.4 47 1968057 12.9 9.", 15.8 16.4 Pv49 481 405 -76 -1412 RDas - T.1 2.1 5.0 -2.9 8.4 -2.5 -5.8 2118 -8.8 -3.0 11.7 48 19484 5 23.6 29.8 24.T 3.9 PV50 405 507 102 1230 W Tc

-9.1 -3.1 12.2 49 237067 2.9 -15.0 3.6 18.7 Pv51 507 395 -112 -1102 eD4f -9.5 3.0 6.5 2208 819 RU3b 10.5 -5.1 -5.4 2208 -9.1 -3.1 12.2 50 1964924 23.8 31.8 22.3 6.8 PVS2 395 646 251 -9.1 -3.1 12.2 57 694340 8.4 - T.8 12.5 16.8 Pv53 646 450 -1 % -652 Roth -9.4 4.8 4.6 2208 18.T 27.4 16.8 3.6 Pv54 450 562 112 1857 ru5d 12.6 -4.0 -8.6 2208 -9.1 -3.1 12.2 52 1542314

-489 RD1d -T.6 3.2 4.3 2208 -9.1 -3.1 12.2 53 672901 8.2 -6.3 19.5 16.5 Pv55 562 448 -114 -9.1 -3.1 12.2 56 1749913 21.2 29.2 19.8 5.2 Pv56 448 619 1 71 1493 M7e 11.2 -4.2 -T.0 2208 901c -8.6 4.1 4.5 2208 -9.1 -3.1 12.2 ST 639273 T.T - T.8 10.9 16.7 PVST 619 471 -148 -678 18.6 23.1 13.2 6.4 Pv58 4 71 563 92 1582 ruTe 8.4 -2.5 -5.8 2208 -9.1 -3.1 12.2 58 1534824 Roff -5.1 1.8 3.4 2208 -9.1 -3.1 12.2 59 313303 3.8 -9.4 3.5 13.6 PV59 563 508 -55 -291

-9.1 -3.1 12.2 60 1265689 15.3 19.2 12.8 T.2 Pv60 508 646 138 438 WU6e 8.7 -3.T -5.0 2208 8.3 -6.0 10.7 16.5 Pv61 646 540 -106 -597 PD1d -T.6 3.2 4.3 2208 -9.1 -3.1 12.2 61 e86954 rWd 5.0 -1.4 -3.6 2208 -9.1 -3.1 12.2 62 1137791 13.8 13.5 13.1 8.6 Pv62 540 602 62 1095

-389 acte -6.4 2.4 3.9 2208 -9.1 -3.1 12.2 63 571848 6.9 -6.T 8.2 16.1 Pv63 en2 518 -84 4.8 T.2 518 650 132 1179 ru6e 8.T -3.7 -5.0 2208 -9.1 -3.1 12.2 65 392579 5.7 -0.7 m Pv64

-9.1 -3.1 12.2 67 525551 6.4 -8.6 8.2 16.5 650 524 -126 -742 cid -T.6 3.2 4.3 2208 k Pv65 Pv66 524 575 . 51 1062 eu6d 5.0 -1.4 -3.6 2208 -9.1 -3.1 12.2 68 1104525 13.4 5.T 13.0 12.6 8.6 6.0 15.6 c.re

-178 aD1f -5.1 1.8 3.4 2208 -9.1 -3.1 12.2 69 471165 -6.9 Pv67 575 520 -55 -9.1 -3.1 12.2 70 1106494 13.4 15.9 11.1 T.4 Pv68 520 640 120 1037 ru66 T.8 -3.0 -4.8 2208 Role -6.4 2.4 3.9 2208 -9.1 -3.1 12.2 71 587294 T.1 -6.5 8.3 16.1 Pv69 640 567 -73 -568

-9.1 -3.1 12.2 T1 582294 T.1 4.9 4.5 8.6 Pv70 567 608 41 794 ru6d 5.0 -1.4 -3.6 2208 meJeIwJn StPESS 50.0 40.4 21.5 Mist u StatSS -26.1 -6.2 -6.3 maurt u MANGE 76.1 46.7 27.8 Notes:

tineerised Stress ($12) = themnt stress + pressure stress

• C2(piping stress), C2 = 1.28 tineerizef Stress $523) = thernet stress + pressure stress + C2tpiping stress), CT = 1.23 tineerired Stress ($31) = themst stress + pressure stress, rio piping stress ee: C2 = 1.28 l

l t

0 0 0 0 1A8tE 6-3: max Mun staESSES, TRAustEuf 1A1 STRESS LINE 10, tus!DE SURFACE, SAFE END-10-EL80W WELD Piping Stresses penisus Stress PV # Strt Erd Detto' Max Therzust Stresses Pressure Stresses pipe panwit Stress (sae notes)

Tempfewp Temp #any Sese 512 523 531 Rete min $12 523 531 Press $12 523 S31 toed Mrt sn) mc/1 (ksi) (ksi) (ksi)__; (psi) (ksi) (ksi) (ksi) pv # (in-Ibs) (ksi) (ksi) (ksi) (ksi)

( F)( F) ( F) (F/Mr) Used _________ === =__________w.w.

0.0 0.0 0 0.0 0.0 0.0 1 24024 0.3 0.7 0.T 0.0 Pv1 70 70 0 0 5s70 0.0 29.8 87.8 55.8 -6.5 aute 14.7 -7.7 -T.0 578 -2.7 -0.6 3.3 2 2457310 Pv2 70 -482 412 848

-2.7 3.3 3 522099 6.3 -48.0 33.8 43.4 Pv3 482 163 -319 -3954 AD2s -39.7 43.1 26.6 578 -0.6 4 30.4 89.3 58.1 - T.4 14.8 - T.2 -T.6 578 -2.7 -0.6 3.3 2508422 Pv4 163 482 319 758 Rt'3e

-2.7 -0.6 3.3 5 280908 3.4 -54.8 27.1 43.4 PV5 482 97 -385 -3849 t02e -39.7 13.1 26.6 578 30.1 82.2 62.6 -6.0 Ruff 10.7 -4.0 -C.T 578 -2.7 -0.6 3.3 $ 2486005 Pv6 97 240 143 692

-2.7 3.3 6 2486005 30.1 52.6 74 2 12.0 PVT 240 104 -136 -292 RD5e -9.0 3.6 5.4 578 -0.6 94.T 57.1 -13.1 18.9 -T.5 -11.4 578 -2.7 -0.6 3.3 6 2486005 30.1 Pv6 104 338 234 1305 au2c -37.0 26.1 33.8 aD3e -30.2 10.1 20.1 578 -2.7 -0.6 3.3 7 411227 5.0 PY9 338 101 -237 -2373 22.2 73.2 41.0 -12.2 au2d 16.9 -6.1 -10.8 578 -2.7 -0.6 3.3 8 1829208 Pv10 101 269' 168 1218 3.3 483188 5.9 -3.3 18.3 12.0 Pv11 269 135 -134 -651 RD5c -9.0 3.6 5.4 578 -2.7 -0.6 9 au2g 9.8 -2.6 -T.2 578 -2.7 -0.6 3.3 9 483188 5.9 24.7 8.9 -6.7 Pv12 135 178 43 1538 3.4 - T.0 11.4 11.2

-724 RD5d -T.T 2.8 4.9 578 -2.7 -0.6 5.3 11 280977 Fv13 178 110 -68 15.0 -4.9 -10.1 578 -2.7 -0.6 3.3 12 2195571 26.6 80.5 53.1 -11.1 Pv14 110 248 138 1390 mu2e

-2.7 3.3 13 429664 5.2 -4.8 16.8 12.0 Pv15 248 114 -134 -376 RD5e -9.0 3.6 5.4 578 -0.6 60.8 28.6 -12.2 16.9 -6.1 -10.8 578 -2.7 -0.6 3.3 14 1382962 16.8 PV16 114 266 152 1717 au2d 4.6 -8.0 16.6 12.4

-666 RD56 -10.1 4.4 5.7 578 -2.7 -0.6 3.3 15 378492 PVIT 266 117 -149 -0.6 3.3 1648571 20.0 58.9 39.2 -6.0 Pv18 117 255 138 873 autt 10.7 -4.0 -6.7 578 -2.7 16 3.6 5.4 578 -2.7 -0.6 3.3 17 44C821 5.3 -4.5 17.1 12.0 Pv1? 255 130 -125 -428 RD5e -9.0 20.5 69.5 37.3 -12.2 ru2d 16.9 -6.1 -10.8 578 -2.T -0.6 3.3 18 1695162 Pv20 130 321 191 1192 3.3 395736 4.8 -8.3 18.0 12.4

-823 RD5e -10.6 4.9 5.7 578 -2.7 -0.6 19 i

to Pv21 Pv22 321 140 -181 140 478 338 1294 Ru2e 21.2 -9.6 -11.6 578 -2.T -0.6 3.3 23 139825 1.7 18.6 32.9 -11.T -13.3 22.1 50.2 13.5

• -352 -11.5 5.1 6.4 578 -2.7 -0.6 3.3 24 1536736 Pv23 478 305 -173 RD1e

-0.6 3.3 1536736 18.6 54.8 38.3 - T.4 ,

305 380 75 1767 au4e 10.3 -2.6 - T.T 578 -2.T 24 Pv24 a05c -9.0 3.6 5.4 578 -2.7 -0.6 3.3 24 1536736 18.6 26.1 47.T 12.0 Pv25 380 253 -127 -395 18.6 54.8 38.3 -T.4 10.3 -2.6 -T.T 578 -2.7 -0.6 3.3 24 1536736 Pv26 253 300 47 1572 au4e 4.1 -2.7 -0.6 3.3 25 306030 3.7 -4.0 11.0 10.1 Pv27 300 231 -69 -139 eD5e -6.2 2.1 578 15.8 53.1 27.2 -T.5 au3e 13.2 -5.5 - T.7 578 -2.7 -0.4 3.3 26 1302842 Pv28 231 431 200 873 578 -2.7 3.3 27 406195 4.9 -9.4 18.6 13.5 Pv29 431 243 -188 -807 RD1c -11.5 5.1 6.4 -0.6 11.3 43.9 19.6 -11.2 14.1 -3.9 -10.2 5 78 -2.T -0.6 3.3 23 936874 Pv30 243 328 85 1466 au4d 578 -2.7 3.3 29 133506 1.6 -11.1 T.3 11.2 Pv31 328 234 -94 -197 RD5d -T.T 2.8 4.9 -0.6 5.2 30.2 1.3 - T.5 884 au36 - 14.1 -6.4 - T. T 578 -2,7 -0.6 3.3 31 428891

( Pv32 234 4TT 243 -2.7 3.3 30 1689142 20.5 24.1 56.6 13.6 4TT 234 -243 -350 RDie -13.0 6.5 6.5 578 -0.6 l Pv33 ausb 20.3 -6.8 -13.5 578 -2.T -0.6 3.3 32 1354047 16.4 65.1 26.7 -16.2 Pv34 234 418 184 1357

-2.7 -0.6 3.3 33 929878 11.3 T.5 29.6 14.7 Pv35 418 344 -74 -14C0 PD49 -10.2 2.9 7.3 578 46.T 29.1 -6.1 l 10.1 -3.3 -6.8 578 -2.7 -0.6 3.3 34 1248528 15.1 Pv36 344 429 85 1241 PU3e

-3.1 -0.7 3.T 35 458860 5.6 -8.4 20.0 14.0 Pv37 429 276 -153 -997 RDie -11.5 5.1 6.4 659 19.2 66.4 35.4 -13.3 Pv38 276 411 135 1751 . Rose 17.5 -5.3 -12.2 708 -3.3 -0.7 t.0 36 1585180

-10.3 4.2 6.1 939 -4.4 -0.9 S.3 37 194453 2.4 -15.5 10.7 15.5 Pv39 411 282 -129 -848 RD1d 25.3 72.3 45.3 -0.8 Pv40 282 5VI 309 1058 203. 14.8 - T.2 - T.6 1543 -T.2 -1.5 8.8 38 2092208

-T.3 -1.6 8.8 39 1335898 16.2 10.6 40.0 23.9 Pv41 591 524 -67 -2433 R039 -12.2 3.1 9.1 1551 21.8 51.9 45.4 3.1 7.1 -1.9 -5.2 1587 - T.4 -1.6 9.0 40 1799577 524 584 60 659 Ru6d Pv42 9.4 A6453 17.5 15.5 44.8 20.4

[ -

584 469 -115 -494 RDid -10.3 4.2 6.1 1655 -T.T -1.7 41 Pv43 auTc 11.8 -3.4 -8.4 1674 - T. 8 -1.T 9.5 42 5372581 22.T 60.5 45.1 -1.0 Pv44 469 546 TT 1406 8.3 -8.3 25.0 21.6

-11.5 5.1 6.4 1765 -8.3 -1.8 10.0 43 685581 Pv45 546 376 -170 -946 RDic m-

g 1 TASLE 6-3: nextMtm sinESSES, TRANSIENT 1A1 -

STRESS LINE 10, INSIDE SueFACE, SAFE END-TO-Els0W WELD PV 8 Strt F.nd cette Men Thernet Stresses Pressure Stresses Piping Stresses monie m Stress Tseptemp Temp amup Base pipe mammt Stress (see sotes) mete nun 512 S23 531 Press $12 s23 S31 tend pr(myr) ac/1 s12 sz3 531

( F)( F) ( F) (F/Mr) Used (ksi) (ksi) (ksi) (pol) (ksi) (ksi) (ksi) pv 8 (in-tbs) (ksi) (ksi) (ksi) (ksi)

Pv46 376 468 92 2209 au5e 14.6 -4.1 -10.5 1778 -8.3 -1.6 10.1 44 1567425 19.0 55.5 35 5 -3.5 b PV47 468 344 -124 -558 acid -10.3 4.2 6.1 1968 -9.2 -2.0 11.2 45 356343 4.3 -16.T 1s . f 22.5 Pv48 344 481 137 1670 auce 17.5 -5.3 -12.2 2020 -9.5 -2.0 11.5 46 1805393 21.9 65.2 30.9 -4.4 Pv49 481 405 -76 -1412 n043 -10.2 2.9 T.3 2062 -9.7 -2.1 11.7 4T 1068057 12.9 3.1 31.7 24.9 Pv50 405 507 102 1230 auTe 11.8 -3.4 -8.4 2118 -9.9 -2.1 12.0 48 TM8405 23.6 60.2 46.T 2.0 Pv51 507 395 -112 -1102 a04f -13.4 4.1 9.3 2208 -10.3 -2.2 12.5 49 237067 2.9 -25.8 10.2 28.9 Pv52 395 646 251 819 nU3b 14.1 -6.4 -T.T 2208 -10.3 -2.2 12.5 50 1964924 23.8 63.9 42.2 3.6 Pv53 646 450 -196 -652 a01b -12.6 6.1 6.5 2208 -10.3 -2.2 12.5 51 694340 8.4 -12.2 26.2 24.8 Pv54 450 562 112 1857 au5d 17.6 -5.4 -12.2 2208 -10.3 -2.2' 12.5 52 1542314 18.T 57.1 32.2 -3.2 Pv55 562 448 -114 -489 acid -10.3 4.2 6.1 2208 -10.3 -2.2 12.5 53 672901 8.2 -9.2 22.6 24.2 Pv56 448 619 171 1493 auto 15.4 -5.4 -10.0 2208 -10.3 -2.2 12.5 56 1749913 21.2 59.7 37.8 0.2 Pv57 619 471 -148 -678 acte -11.5 5.1 6.4 2208 -10.3 -2.2 12.i 57 639273 T.T -12.1 23.2 24.6 Pv58 471 563 92 1582 auTe 11.8 -3.4 -d.4 2208 -10.3 '-2.2 12.5 58 1534824 18.6 43.1 35.0 2.6 Pv59 563 508 -55 -291 a01f- - T.2 2.4 4.8 2208 -10.3 -2.2 12.5 59 313303 3.8 -14.4 9.8 22.2 Pv60 508 646 138 438 au6e 11.9 -4.8 -T.1 2208 -10.3 -2.2 12.5 60 1265689 15.3 40.9 25.3 4.5 Pv51 540 -106 -597 RD1d -10.3 4.2 6.1 2208 -10.3 -2.2 12.5 61 686954 8.3 -8.9 23.0 24.2 Pv62 646 540 ~ 602 62 1995 au6d 7.1 -1.9 -5.2 2208 -10.3 -2.2 12.5 62 1137791 13.8 30.0 26.2 T.3 Pv63 602 518 -84 -389 acte -8.8 3.2 5.6 2208 -10.3 -2.2 12.5 63 571848 6.9 -9.T 18.2 23.4 m Pv64 518 650 132 1179 au6e 11.9 -4.8 - T.1 2208 -10.3 -2.2 12.5 65 392579 4.8 14.5 0.9 4.5 650 524 -126 - 742 a01d -10.3 4.2 6.1 2208 -10.3 -2.2 12.5 67 525551 4.4 -13.4 18.5 24.2 h

u Pv65 Pv66 524 575 51 1062 au6d T.1 -1.9. -5.2 2208 -10.3 -2.2 *2.5 68 1106525 13.4 29.0 25.3 T.3 Pv67 575 520 -55 -178 aoff -T.2 2.4 4.8 2208 -10.3 -2.2 12.5 ~69. 471165 5.7 -10.0 14.2 22.2 Pv2 520 640 120 1037 ausb 10.7 -3.8 -6.9 2208 -10.3 -2.2 12.5 70 11084M 13.4 34.7 22.4 4.8 Pv69 640 . 567 -73 -568 nele -8.8 3.2 5.6 2298 -10.3 -2.2 12.5 T1 582294 T.1 -9.4 18.5 23.4 Pv70 567 608 41 794 au6d 7.1 -1.9 -5.2 2208 -10.3 -2.2 12.5 71 5822M T.1 i 14.4 18.7 T.3 Nextuun staESS M.T T4.2 43.4 minigam StaESS -54.8 -11.7 -16.2 MnalMUM RANGE 149.5 85.9 59.6 Notes:

mesimus Stress ($12) = som therset stress + (K3-1)(tineer thernet stress) + K1(pressure stress) + C2(K2)(pipine stress) per.isua stress (523) = men thernet stress + (Ki-1)(tineer thernet stress) + K1(pressure stress) + C2(K2)(piping stress)

Maximun Stress (531) = ses thernet stress + (K3-1)(tineer thernet stress) + K1(gressure stress), no piping stress dere: K1 = 1.2, K2 = 1.8, K3 = 1.7, C2 = 1.28 I

I l

.. 3 O O TAetE 6-4 TMERMALSTRESSBASECASESFORTHEN0Z2LE-TO-SURGELINENELD RAMP T START T END T HL ----~ INSIDE SURFACE STRESSES (PSI) -- - -- ----0UTS10E SURFACE STRESSES (PST)-~~----

RATE

= LINEAR I2ED------- ----SURFACE-----

(F/NR) (F) (F) (F) S12 S23 531 S2 $3 S12 S23 531 S3

............................... ......-..-.--LluEARIZED------

....- ---. SURFACE-----

....a- - --

S2 10000 132 361 108 22723 -4214 -18510 -26737 -24593 -2137T 4861 16516 19198 11686 3500 130 2M 80 11618 -2360 -9259 -13332 -11909 -1073* 3585 7346 10410 $341 3500 130 285 80 13465 -2738 -10728 -15353 -13703 -12663 41 % 8532 1210T 6239 3500 130 320 80 15829 -3271 -12558 -17882 -15882 -14892 4924 9968 14324 7344 1300 130 260 80 9811 -2333 -7478 -10770 -9102 -9231 4037 5194 9317 3920 1300 130 285 80 10847 -2666 -8182 -11856 -9910 -10206 4613 5593 10383 4246 1300 130 300 80 11429 -2864 -8565 -12443 -19326 -1C753 4956 5797 10993 4420

-1000 265 145 80 -2413 -28 2441 3222 3790 2269 1931 -4200 -1220 -3029

-1000 265 125 80 -3145 223 2921 3984 4327 2957 149 -4427 -1992 -3223

-2000 280 130 80 -5444 339 5105 6886 7433 5120 1773 -6893 -3615 -4963

-2000 280 100 80 -6890 728 6161 8438 8682 6480 1102 -7582 -5039 -5486

-3000 - 300 175 80 -4441 -58 4499 6063 7022 4176 2846 -7022 -2336 -4943

-3000 300 150 80 -6103 279 5824 7905 8660 5740 2322 -8062 -3862 -5725 '

-3000 300 115 80 -8207 739 7468 50223 10678 7719 1577 -9296 -5826 -6659

-10000 430 210 140 -13015 1203 11812 16541 17315 12242 3783 -16025 -8504 -11284 2500 250 345 250 11757 -3111 -8646 -12935 -10498 -11062 5636 5426 11381 4078 2500 250 410 250 16032 -4045 -11988 -17545 -14519 -15085 6983 8102 15348 6119 2500 250 480 250 19690 -5046 -14644 -21385 -17611 -18526 8499 10027 18921 7640 1500 200 300 200 10379 -2695 -7684 -11329 -9246 -9766 4874 4892 1006t 3712 1500 200 530 200 18763 -5579 -13185 -19880 -15332 -17655 9603 8052 18799 6379 1500 200 600 200 20189 -6246 -13963 -21280 -16088 -18997 10740 8257 20442 6624 m -700 360 280 250 2745 -16T4 -1072 -2289 -292 -2584 4876 -2292 4028 -1503 h

co

-aco

-1500 360 410 250 315 250 255 1669 1943

-1308

-1755

-361

-187

-1162

-1153 521 1116

-1572

-1830 4261 5603

-2689

-3774 2912 3703

-1836

-2530

-1500 410 255 255 -1048 -9 71 2019 2072 3766 984 4382 -5365 816 -3769

-3000 495 310 250 -3290 -835 4125 4990 6943 3093 5297- -8390 -565 -5900

-3000 '495 250 250 -6492 . -90 6581 8460 9906 -- 6105 4147 -19252 -3586 -7327

-10000 495 250 250 -13620 1216 12404 17265 18187 12811 4429 -17240 -8796 -12194 i 2500 300 400 300 14295 -3840 -19455 -15592 -12551 -13450 6978 6471 1392T 4902 2500 300 500 300 20053 -5263 -14789 -21666 -17641 -18867 9114 9754 19463 7475 2500 300 600 300 23796 -6538 -17258 -25467 -20391 -22390 11018 11372 23323 8833 1500 350 450 350 14559 -4226 -10333 -15512 -11977 -13699 8080 5619 14647 4389 1500 350 500 350 16723 -4862 -11862 -1TT75 -13752 - -15735 9054 6682 16804 5243 1500 350 600 350 19647 -5951 -13696 -20T38 -15763 -18486 10765 7721 19904 6151

-1000 440 350 350 2753 -1961 -792 -2112 326 2592 5980 -3388 4434 -2271

-2000 560 400 400 -290 -1480 1770 1503 3901 270 6429 -6699 2137 -4705

-3000 480 350 350 -1098 -1355 2454 2584 4875 1031 5932 -6963 1437 -4817

-10000 550 370 330 -6855 -212 7067 9653 11461 6446 6280 -12726 -IT88 -8878 1500 500 600 500 19132 -5814 -13318 -20254 -15281 -18002 10941 7061 19458 5594 1500 500 650 500 20646 -6410 -14236 -21791 -16291 -19427 11777 7650 21064 6112 1500 500 680 500 21310 -6743 -16567 -22448 -16635 -20051 12187 7F45 - 21795 6318

-700 615 *60 550 - 7080 -3206 -3874 -6647 -3000 -6664 9097 4 433 8980 -1516

-1500 625 530 530 2857 -2235 -622 -1869 1101 -2691 8212 -5521 524T -3822 -!

-2200 650 530 530 -82 -1614 1696 1521 4174 . 74 7864 -7938 2757 -5609 0 70 70 70 0 0 0 0 0 0 0 0 0 0 0 615 615 550 11750 -4299 -7452 -11843 -7593 -11058 VJ299 759 13210 889

g S

w TABLE .) ]

STRESSES AND FATIGUE USAGE FOR A TYPICAL PV AT THE N0ZZLE-TD-SURE LINE WELD SECit04 20. 9 AT TWETA = 39. DEGREES FOR S12 05 TME IWSIDE SURFACE Pvf PVW TMAX TMIN 3*SM SY E ALPMA SW SP KE SA MACTUAL matt 0W U UStM (F) (F) (PSI) KPSI) (PSI) (IN/IN/F) (PSI) (PSI) (PSI) 22 46 550. 200. 56250. 19350. 2555E+08 .9500E-05 64711. 11M15. 1.501 9M60. 13. .19012E+04 .0068 .0008 Pt.

• Pt

• O - TNERWil SEleIWC = 39992. PSI ALLME TieERNAL STRESS RANGE AS PERMITTED SY RATCMETING RCLES = 193272. PSI TRAirS. Pvt N CIMPUTER Aun am4E -------!WSitFE SUkFACE ST# ESSES (PSI) --------OtfTSit4 SURFACE STWESSES (PSI) -

NM4E -----LINEARIZED----- - SURF ACE--~~-- ----LluEARIZED ------ SUR F ACE ---

$12 523 $31 S1* S2* S3* S12 S23 $31 $1* S2* S3+

i MU1A1 22 FX GOtY . of.

'.'.. -59. O. -742. -57. 803. -713. -93. O. -960. -116.

FJ1A1 22 P9ESS Gouw -3778. -4830. 8609. -1527. s.C 7119. -4642. -2029. 6671. O. 5627. 621 MU1A1 22 MT G0JU 7688. -6985. -703. O. -7766. -645. 11248. -1555. O. -12977. -tT89.

o MU1A1 22 STRAT GTAE GGw 14357. -13736. -621. O. -13546. 1443. 15360. -12364 $997 9. -17583. -1509.

• 1 B i.

to MUTA1 22 TRANS GOUG GILA 23796. -4538. -17258. O. -25 4 7. -20391. -22390. '41018. 113'2. O. M. 8833.

MU1A1 22 TOTAL 42802. -3276P. -10e32. 77539. 66535. -118G4 380. -13782. 13401 . 2254. -145 D ,, 12329. ,

r TRANS. Pvf LOmo CIMPUTER RUN NNUE --------tWSIDE SURFACE STRESSE3 (PSI)------ ------6STDE SURFAQ STRESSES (PSI) 1 mmE - '--Lt-EARIZZD------ --

--SURFACE-*-* - -- LisEARIZES---- ~~~-SURFACE

• S12 523 S31 S1* $2* S3* 312  ?.23 IB31 S1* S2e 33*

t' CD181 46 Fx GCwf 33.- -35. -3. O. -36. -3. 41. -36.. -3. O. -49 -6.

CDist 46 PRESS GGW -1MT- -T338.- 2385. -423. 593. 1972. -1286. +562. 1848. P. 1559 1986.

CD181 ' 46 942 G0JU 1904 -912. -92. C. -1014. -84. 1469. -1266. -203. 3. -1695. -230.

CD181 46 STRAT 00EY GEMY -TTG6. 2471. 5234 6. 6446. 4737. 7386. -21M. -5207 O. -TTIS. -4650.

CD1*1 44 TR*Je3 E08J EOEJ -14196. 1312. 12886. O. 18045. 18889, 13355. 4127. -17' 4 . O. -9277. -12310.

(D?81 46 70TAL -21908. 1497. 20419. -40876. -14942. 25934 20965. 72. '.f0E3. 18742. -3453. -73209. ,

• ieciE, FOR TOTAL LOC, TNES2 REPRESEWT STRESS DIFFERERCES $12, 523, "A S31.

e T

4 r

• i y -

., y F

4 - -

c-  :#

e 4

d O

FIGURE 6-1 i CEOMETRY OF PREG URIZEA SURGE N0ZZLE l

44.625" IR j 49.375" OR .

\' '

y 4.375" i

i 4,625"

=

/ 7.625" 4

/

< ,, a l ::

8 s.

N-c

1. h

, ws l

l

.. m5 I

- Sa75-

-O . , -

4I i

l. 6-30 I~ '

_.m_.....

,% FIGURL f-2 FINITE ELtMENT MODEL OF FRESSUR1212 JURGE N0ZZLE

\/

/\/ .

V

1. \/N/%

' N / 9, /

\, N/\/N/

/\/N/N/ cl a ddi ng

/N/N/N/

/N/N/\ /

/N/N/ '% / TA

/N/i~I,/T/% ' ['< ',-

s I S: o, 1o- Q' ,.... .<

. , ,7 h

. *I k e, o

I o ., _~. 3

' .r.

prt1.$ crilet' l r.er head .Q[)

l ,' '. s ' g J

, ~, . "a . ,i,,,,.

...., . ,1.;,,,*.),,< s, I

i i'.h *. * ~. ,

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i..

l i

'. i.8.e. '.2. i..

'344..

W

.t). .

!.'l.'.*..

Q tafe end-to-nozzle weld $5E

.:::.,. .a.

safe end -

,9.7

..::}

sa 'e end-to-elbow weld elbow  ::,

p 9 4'

]'

6 31

x , ' . b i

O 0 8

l r D g N , r

, E g ze G opr E gTT L -

B G

C r -

R ts

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pr l

e (

z - /

T z N

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-i /

I S

N \

[l N

A r I Y R z  !

j J )

r 3 T P b 6

- L A ,

_ i

. (

C \, ~

0 E I r '

4 z e OUR G

I F

Y T P

_ \

, \. i m

F F O -

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T E

L P

s e

%n M r A u _ .

X e E t )-

a r

e p

_ n

[l i

l m

T e b 0 2

5

_ q

_ \

1 I 7 I

/

l (

0 0

8 e

7 -

0 0

5 4

g e

3 e

0 2

0 0

1

~ r. ' 3* e W" ,

v13 J

FIGURE 6-4 k

SUMMARY

OF THERMAL BOUNDARY CONDITIONS (pressurizer surge nozzle)

\

D /\/N

\/\/N

\/\/N/N

1. \/N/N/N region 1 ,

/N/N/N/N

/N/N/N/

/N/N/N/s j fN/N/  % / N' ,--

/N/*//N s //\-.Tl'/_

4-i1 region 2 f ' '

insulated / '~ v <

Tl'$,'$;o o-

' , a, 'i,t

., .,i,"

(]

s s.

s,:.::.b,.

IIalIeai 4l441li[ **

u:

.l.;.l.l.l * * '.l.4'[

.cf o;: region 3 i

NOTE: Regions 1 to 5 have bulk fluid Tiiiri temperatures and film coefficients I'E applied. r"$'.

7m..

v.. .

f Tim

07: . region 4 insulated -l-l :.l-IO:

.#)

'.,$5.).

I region 5 db E

.C

$L i 2 insulated (symmetry) 6-33

FIGURE 6-5 IDCATION OF DELTA-T VALUES O- (pressurizer surge nozzle)

N

\/N

/\/N

/\/N/x

\/\-/N/N

1. \/\/N/N

/\/-N/N/N h

/N/N/N 's

/ N / N-

,s s/',/'i.s

~'

s ~h _ @

i s /t ','hllS:'h,:'~h'?,.:;

N - .'

- 3 S', <i

/ N' , i,i' <i

,~  ;,~(15

. .e:;, ,, i

-o ,.:.

is.+' R D h Delta-T Locations

@:!:!:!:%!-ji,

'-t '.-j, 1 to 9 are radial Delta-T h) 10 to 15 am axial Delta-T g :.h: n; h

f. ~I h-

'~"

e@ .

O .

~

6-34

i FICURE 6-6 STRUCTURAL BOUNDARY CONDITIONS (pressurizer surge nozzle)

{

l 1)

\/N

/\/N

' \ / 'N '/ N

' % / N ' t 'm /

/ \ / N / N /. N 2)

/ N / N:/ N / N

/N/N-/N/^

/N/%l l

/ %. I % / % /l T

/N/)-l%

% ~

,o i s I st ,I * *. -

l . sis

.,/ g.

,I

~ /0/ ,- 3t .M,'jd:l

~...,,,,.

h,,

'"j 3)l'( ,

w:e:v.u

1) Nodes restrained from motion .w. .U.'.'.

nomal- .to' surface tl-l-l-l-l-l-l-

. iiiiiiii

2) Pressure applied to internal surfaces. c. .v0j;

~;~fj'ef...

3) Temperatures applied at each node. $l- 2).

7iW iiiiii hk.

u.

,3.'.'.

.T.f.;.h. -

. .i.'.i.

3)

$..[

....a O

' nodes coupled _ in axial direction.-I

~

for symmetry 'h end cap pressure 6-35

FIGURE 6-7 IhCATION OF STRESS CLASSIFICATION LINES (pressurizer surge nozzle)

N

\-/ x

'\/N g'\/N/x

'\/\/N/x @

1. \/N/N N

/\/N

/N N/N N/N/ N  %

N / N. /N / s' /T\ , - 3

/s/ /  % / 9 ,, ... <, i.C , e, e i% / % / N~ ,, C p-(.e e<.,,,,,;- ~., ',s c

/ Ss ,, ~ <. ;t< o - i s, < -

^

~

• ~,C l1l',*.iO,',

i~, i U4 l

1 O .W:i.i:

li44t8i1

@ sct nunber i.'M

c. . .

'2W--

m M 8

.. )!

A'.;. .

sery M 10 T.*

9 6*

m 4n

$i M. .

ee M

O 6-36

O FIGURE 6-8

-GEOMETRY OF HOT LEG SURGE N0ZZLE-Ho t leg 58.6" IR 61.725" OR O

1 Cla dding .

Surge Line-To-Nozzle Wel d

{

o

_[

Surge Line -

-n g a

/ Surge Nozzle:

o i

5.375" -5.75" 8.625" 4 .375" l l l i

Center Line Of Nozzle ,

4 4

4 6-37

__, , , , . , , , . . , , ~ - - - _. . . . - . . . -

,,a

a

) FIGUPI 6-9 FINITE ELEMENT K) DEL OF HOT LEG SURGE N0ZZLE  !

4

/

O O

! k

/ 1 \ \

8 i \\

E!liiRillfgifgMglF i EB~ \\\s b

OL 6-38

FIGU 0

- EXCPLE OF 'A TYPICAL TRANSIENT FOR-THE HOT LEG SURGE N0ZZLE Temperatures - Pzr, HL Nozz1e (top / bottom), RC 7m 7 ,

/

l F- 7- '

y l 7

sm j M '

saa j

y g _ .. _ .. y .. _

m

/ $v

/

, 4.  !

~ i .

I 3

% / )  ?

T -

w c j t  : .

$~ I' Ah W 200

/ ,

/ .

}O //

/ y

+t W ,,".

'_ ~~~~' -

100 , , , , - 7,p If ~

_ _ gotto.

- rpzr Trc 0

0 20 40 60 00 Time Chr)

FIGURE 6-11 O* I4 CATION OF DELTA-T VALUES (hot leg surge nozzle) h- DELTA-T LOCATIONS 1 to 10 are radial Delta-t 11 to 15 a re a xial Del ta-t O

O g -

@1 r

=

@/

=, :

in

, 1w fginu=-Q jfE. [

\\\s _

1

\()-

6-40

-.-.r.

.4

.g FIGURE 6 O_:-

HOT LEG SURGE N0ZZLE TEMPERATURE CONT 0URS (F) FOR A TYPICAL _PV -

• s

?

?

..p r'

8M- -g lLO i.

.- 120 ---

160' 200' 24 0 ,

320 2@ I

p. < KK _

r 1

-1 t

~

[

p:

L-1

.g

!- 6-41 4

--4 -, . - - , . _ - e .-.......m.,_ _ , .,,,,...m_,,v,..._,,,...w .. ,,.y ,4.-,.,,, , .. _,..r,r,-,,_..,,, -

i i

FIGURE 6-13 thCATION OF STRESS 01ASSITICATION LINES (hot leg surge nozzle) t 9

0 O

rarsnuman d!; -

@ GO @ @ O @

O 6-42

)

("j

,t T

7.

SUMMARY

OF RESULTS The B&W Owners Group has developed a program to comprehensively address the requirements of NRC Bulletin 88-11, " Pressurizer Surge Line Thermal Stratification". The Owners have collected the necessary information required to evaluate the surge line. In addition to operational records and plant design information,. plant thermal stratification data and thermal striping test data have been obtained.

It has been determined that the lowered loop plant configuration and plant operations are sufficiently similar for a generic development of the design basis transients. Stratification data for the lowered loop plants has been obtained by monitoring the thermal conditions of the Oconee Unit 1 surge line during

,- normal operations. This data made it obvious that the original design basis did V not adequately represent surge line thermal conditions. The details of the Owners plan have been subsequently developed to correlate the Oconee Unit 1 stratification data and other information for use in generating a set of transients accurately representing the surge line conditions for the operating conditions in the plants.

Davis-Besse Unit 1 (DB-1), a raised loop plant, requires its own instrumentation and a separate plant specific set of design basis transients because of inherent differences in its design. These differences are discussed in Section 3 and the DB-1 analysis will be addressed in a supplement to this report.

Revised surge line design basis transients accounting for plant evolutions affecting the surge line for the 40 year design life of each B&W plant have been developed. The number of occurrences for each transient varies among the plants due te differences in operational history. The plant heatup and cooldown transients have been the most significant contributor to the fatigue usage factor for surge line components.

O U 7-1

p A structural loading analysis of the surge line has been performed to take into V account the global effects due to thermal stratification. The resulting internal forces and moments have been applied for the fatigue stress analysis of the surge line and the associated nozzles.

The fatigue stress analysis takes into consideration the stress ranges for the global effects due to thermal stratification, the localized effects due to thermal stratification, the pressure ranges, the Operating Basis Earthquake, the thermal striping and the fluic flow conditions. All resulting stress intensities have been shown to be within their allowable limits. As a result of the fatigue analyses, the cumulative usage factor is less than 1.0 at all locations of the surge line and its nozzles.

In summary, the following is a tabulation of the highest usage factor of all six lowered loop plants for the most important surge line components.

Surge Line Component Usage Factor (40 year Life)

Surge Line Elbow 0.82

) Straight Pipe Section 0.48 Drain Nozzle Branch Connection 0.37 Pressurizer Nozzle 0.41 Hot Leg Nozzle 0.62 In view of the conservatism accumulated in the synthesis of the design transients and in the analysis of resultant stresses, these fatigue usage values provide assurance that the 40 year licensed life of the lowered loop B&W plants will be I met with c.onsiderable margin to accommodate normal variations in operations. l l

l l

l l

nU 7-2

_ i i

o r

\

8. ' BASES FOR THE B&WOG ANALYSIS - FOR PLANT SPECIFIC-APPLICATION ,

The generation of the revised . Design Basis transients.-and the thermal stratification fatigue stress analysis of the surge line have been-based on conditions stated in this section.

The thermal stratification-fatigue stress analysis has been based on:

. no interference of the surge line with any other structure,

. surge line movement within the travel range of each snubber,

. surge line movement within the travel range of each hanger, and:

. branch moments at- the surge line- drain nozzle connection within their respective. maximum allowables (for Deadweight, Operating Basis Earthquake and-thermal stratification),

The generation of the revised Design Basis transients (for. future events)'has h been based'on the incorporation of operational guidelines which:'

. limit the pressurizer to RCS temperature difference-~during plant heatups and cooldowns (imposed with pressure / temperature limits),

and i

- prevent surveillance -tests s -+ cause rapid additions of water to the RCS from being performed with a pressurizer to RCS temperature difference greater-than 220F. ,

The heatup and cooldown Design Basis transients which were defined for future-operation should remain conservative if the pressure is limited in accordance with Figure-8-1. The curve shown in Figure 8-1 is a composite of variable subcooling limits that change with RCS temperature. The operating procedures.of '

each of the lowered loop plants are to maintain pressure and temperature during heatup.and cooldown operations to the .right 'of the selected maximum allowed subcooling limits.

(v g.3

• r- - - , - ,- ,

- - . . - . . - . - . - - - - . . - - - . . . - . ~ _ - . . - . . - -. ... - -

To meet the pressure limit specified for heatup in the temperature range '/0F to l

150F, preheating the RCS has been recommended. This may be accomplished by throttling back on the decay heat system cooling water (i.e. component cooling water) and/or bypassing reactor coolant flow around the decay heat removal heat exchanger. The availability of decay heat and the requirements of the heatup schedule will dictate the capability of maintaining the recommended P/T profile prior to achieving the conditions necessary for starting an RC pump. The fatigue l evaluation has been performed on the basis that 85% of the heatups for the remainder of the plant life can meet the recommended limit shown by path CDEN in figure 8-1. For those heatups involving pressurization at an RC temperature of 70F to 120F, a less restrictive limit is Included in order to permit RC pump operation at lower RCS temperatures (path ABEN in Figure 8-1) when core decay heat is not adequate for raising RC temperature. The fatigue evaluation has been performed on the basis that 15% of the heatups- for the remainder of the plant life will follow this heatup path. In summary, future heatups have been divided-into path CDEN (85%) and path ABEN (15%).

l O

l O 8-2

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- _ _ _ . _ . _ . . . . - . . . ~ . - . . _ _ . _ . _ . . _. _.

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a Figute 81. Surge Line Operational' Limit- . . j

. j.

2400-t

$ NOT Umit .

l .

.I 2000 - .b .f-e

?. i-( l 1' 1800 - 'M-r/. - -;

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-1600 -

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-!  !! i E 1400 -

! / .i r i Press i Templ SCM -

5.

ei [,, / { A !- 232 ! 50.0 l 350 ! .  !

~/ ' B 'I . 325 t 79.0 350[- ,

h 1200 -

l-E C- s 150 i- 06.0 300 l E -[ 0: 232 100.0 300-r - g' - E ~325 129.0 -300-k 1000

[

F- 325- 179.0 250

/

./_ / G 413- 200.0- 250

,/ _1H 526 225.0 'l 250 800 -

/ 4

/ J l - 526 275.0 1 200

/ / J 728 310.0 l 200 < .l

! 600 - ./ K L 1,000 i 346.5 i 200 - l {

ygs.-.j/ L1 1,400 ~ 389.0 ! 200 f ~ .;

M ! 1,800 422.0 1 200 f'

400 -

.Hentup

/~/ '/ ,G N { 2.200 . 450.51 200

,' B "~

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- 200 -

p C' Surge line limM applicable to j

- both heatup and cooldown. !

I ' ' ' '  !  !'

0 O 200 400= 600 RC Hotleg Temperature, F '

- - Proposed Umit Typical Appendix G Umit iu

~

8-3

O 9.- REFERENCES

1) NRC Bulletin 88-11. " Pressurizer Surge Line Thermal _ Stratification", dated Dec. 20, 1988.

2) BAW-2085, " Submittal in Response to NRC _ Bulletin 88-11, Pressurizer Surge L Line Thermal Stratification", dated May 1989.

3) NRC Letter dated May 18,1990, J.T. Larkins to M.A. Haghi, " Evaluation of__

Babcock & Wilcox Owners Group Bounding Analysis Regarding NRC Bulletin 88-11".

4) H.O. Fuchs et al, " Shortcuts In Cumulative Damago Analysis" in Faticue Under Co'BQ1 L Mading, R.M. Wetzel (ed.), SocAutoEng, 1977.

5) H.Watanabe, " Boiling Water Reactor Feedwater Nozzle /Sparger Final Report,"NEDO 21821-A (February 80).

6) W.S. Woodward, " Fatigue of LMFBR Piping due to Flow Stratification", ASME Pressure Vessel and Piping Canference, Portland, Paper 83-PVP-59, CONF-830607-27 (June 83).

7) T.Fujimoto, K. Swada, K. Uragami, A. Tsuge, and K. Hanzawa, " Experimental Study of Striping at the Interface of Thermal Stratification", Ihermal Hydraulics in Nuclear Technoloav, K.H. Sun et al (ed), ASME (81).

8) T.M. Kuzay and K.E. Kasza, " Thermal Oscillations Downstream of an Elbow in Stratified Pipe Flow", FBR Thermal Hydraulics, Trans ANS, 41, 794-796 (June 84).

9) T.M. Kuzay and K.E. Kasza, " Thermal Striping Downstream of a Horizontal j

Elbow Under Thermally Stratified Flow Conditions", Joint Mta. ANS and AIF.

Washinaton, CONF-841105-10(Nov.84).

l 9-1

-w -- -mi,r.r9 e

10) T.M. .

Kuzay and K.E, Kasza. L" Resolution of - Therma's Striping; _ Issue

\. - Downstream of a Horizontal Pipe' Elbow in Stratified Pipe Flow,'" ANS Annual Meetina. Boston, CONF-850610-4 (June 85).

11) L.: Wolf and U. .Schygulla, "Experimenta11 Results : of- HOR-TEMR ; Thermal l

Stratification -Test in Horizontal _ Feedwater Lines," Trans 9th' SMIRT, Q, - _

pp.361-367.(87).

12)- L. Wolf, U. Schygulla, M. Geiss and E. Hansjosten,." Thermal Stratification 1 ,

Test'in Horizontal Feedwater Pipelines," Proc 15th'WRSRIM,) NUREG/CP-0091, -

L pp._.437-464-_(Feb 88).

~

L '. Wolf. - U. :Schygulla W Geisshand .E. Hansjosten, and A'.- zTalja,

13) .

" Temperature and ~_ Wall- Strain Fluctuations J During- the_- TEMR Thermal .-

Stratification: Tests AT HDR,"-10th=SMIRT, Anaheim,1CA-(Aug.89). -

14) L. Wolf, M. Geiss', E. Hansjosten, and_ A. Talja,'" Evaluation of the Thermal' Stratification Tests at the HOR _ Facility," . NURETH-4' Conference, 2, Karlsruh'e,FRG(Oct89).

4

15) A. Talja.and E. Hansjosten, "Resul. t s of Thermal Stratification: Tests In a-Horizontal < Pipe Line.At the HOR Facil'ity," TRANS 10th SMIRT,t E(89).,

j:- 16)- U. -Schygulla and: L.J Wolf, . -" Design Report - Thermal:' Stratification- -

1 L

Experiments In ' a Horizontal Piping- Section," ' HOR-Test ~ group TEMR T33, l

Battelle-FRG (November, -.1986) . -

l l 17) "ANSYS" Computer Code, ~ Versions 4.lc and 4.3. Engineering Analysis System,_

i User's= Manual Volumes l' and:II,_ Swanson Analysis Systems, 'Inc.

18) "AS'ME Boiler and Pressure Vessel Code", Section III,1986 Edition with' no- -3 Addenda.

.19 ) '

ABAOUS" Computer Code, - Version .4.8.5, 1989, : User's . Manual, Hibbitt, -

- Karlsson and Sorensen, .Inc.

20) J.B. Truitt and P.P. Raju, "Three Dimensional Versus Axisymmetric Finite- )

Element Analysis of a-Cyclindrical, Vessel Inlet Nozzle Subject to Internal-Pressure - A Comparative Study", ASME Pressure Vessel and Piping. Division, Paper No. 78-PVP-6.

9-2

I i

O

10. DOCUMENT SI9 NATURES

'This' document has been prepared by:

A 0. kWbum I ? // 9'/ O M.T Matphews

9. L ' 17 /I A S o,_

R: ' Gurd( ' -s / /

bosta- ILll6lfD ' '

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• . 12./15l96 G.L. Weatherly j The report has.been reviewed for technical content and accuracy by:

u A verp JlR. Gldudehans, Analysis Services g.F %d tzigho

- J/.F. Sheparc ___

Material & Structural Analysis Verification of independent review.

N C.W. Tally, M b- er

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Performance A sis

1. Au K.E. M6 ore, Manager-Materials & Structural' Analysis This document has been-approved for release.

2)

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V 10-1

i Oi 1

1 APPENDIX A-

-A=. Surge Line. Data Acquisition at Oconee' Unit -

O CONTENTS Page

1. Thermocouple Fabrication and Instrumentation Qualification .... A ~2. Data Acquisition System Description and Operation =. .-.-. ... . . .: .

A-3

3. General Description of Data . . . . . . - . ' . - .. . ._ ..- .. . .-. . : . -.. . - .

A-4 1 A-1

__ ' ' . _.-._1-.-- - _-l-- -

~ __. _ ._. . . _ ._ _ .

Oconee Unit 3 was instrumented as part of the B&WOG program as discussed in

'O Section 4.1. This appendix supplements Section 4.1 with additional detail on the data acquisition.

1. Thermocouple Fabrication and Instrumentation Qualification The thermocouple assemblies and associated extension wire assemblies required for instrumenting Oconee Unit 1, including spares, were fabricated by the Babcock &

Wilcox Alliance Research Center (ARC). Seventy thermocouple - assemblies were fabricated from ANSI Type K, 20 gage solid Chromel-Alumel commercial grade assembly wire having parallel conductors individually insulated with ceramic fiber braid, an overall Jacket of ceramic fiber braid and Inconel protective overbraid. The hot junction, or bead of each thermocouple was formed by heliarc welding (fusing) 'the two conductors together, then swaging the bead to a flat disk to ensure intimate contact with the surge line pipe. Standard 2-pole connector plugs with integral cable clamp were attached to the ends of the thermocouple wires opposite the hot junction. Ninety-five thermocouple extension cable assemblies were fabricated from ANSI Type K, 20-gage solid chromel-Alumel commercial grade extension wire having twisted connectors individually insulated I with Teflon, Mylar-backed aluminum foil shielding with drain wire, and an overall extruded Teflon insulation jacket.

Qualifying the commercial grade thermocouples fabricated for the safety-related pressurizer surge line temperature measurement application was accomplished by the standard practice of " type" testing. In this approach, six additional thermocouple assemblies prepared from the same materials and following the same procedures that applied to fabricating the thermocouples installed at Oconee were placed in ovens along with certified Platinum Resistance Thermometers (RTDs) and heated. Comparison between the temperature registered by these qualification l test thermocouples and the reference temperature monitored by the RTDs provided a means of qualifying the surge line thermocouples. Also, as part of the qualification test, the string potentiometers and the linear variable differential transformers (LVDTs) were connected-to the assigned channels of the data acquisition system instrumentation cabinet, as configured for the surge line field test, and operated. Transducers were arbitrarily chosen (four LVDTs and four pts) and exposed to simulated ambient temperatures expected to exist at Oconee Unit 1 during heatup. At these various test conditions, reference

-V A-2 9

-s, ----

b 4

73 displacements were imposed to assess the functional performance of the transducer ,

() types.

Comparison of thermocouple data and the RTD reference indications demonstrated that thermocouple readings were consistently within 1.5F of the reference temperature. This agreement was well within the established acceptance criteria at all test conditions for qualifying surge line thermocouples. From the testing and evaluations of the two types of displacement transducers, it was evident that both the LVDTs and string potentiometers are well suited for the surge line test application. The qualification results demonstrated acceptable displacement transducer operating and performance characteristics when' subjected to the range of expected surge line and pressurizer room ambient temperatures. Finally, the qualification testing process demonstrated the adequacy of the Helios-based data acquisition system for the ambient temperatures expected during plant operations.

2. Data Acauisition System Descriotion and Ooeration The schematic in Figure A-1 depicts the general interface of components which make up the data acquisition system. The system included a Fluke Helios maintame controlled by a hosc Compaq computer utilizing LabTech Notebook V software which was configured to receive the desired data. The computer and Helios front end, which were located in the Oconee 1 control room, interfaced with a remote. Helios extender chassis housed in an instrumentation cabinet located in the reactor containment building near the pressurizer surge line.

This instrument cabinet also contained the power supply, signal conditioning and other interfacing equipment required for the surge line therraocouples and displacement transducers.

With the system installation complete, the integrity of each instrument and acquisition component was checked. Proper electrical loop resistances for the q thermocouples, lead wires, and extension cables were verified. A polarity test and a complete checkout were then performed for all instrument channels. Data collection was started one hour before the pressurizer heaters were turned on and continued until the plant had reached full power and remained there for several days. There were only short time periods in which data collection was interrupted in order to download the data from the host computer, p

V A-3 1

c, .

l .-

fm In addition to overseeing Data Acquisition System operation, members of the surge b line data acquisition team logged important plant events as the control room operators and plant computer system kept the acquisition team informed of the heatup procedural steps being taken. After the heatup was complete, the control room logbooks and procedural checklist were also obtained to assist in associating plant operations to surge line transients.

3. General Description of Data The temperature and displacement of the surge line were monitored, as well as the reactor coolant system (RCS) conditions. A list of recorded parameters is contained in Table A-1. To monitor the thermal conditions of the surge line, thermocouples were spot welded to the outer diameter of the pipe at nine different axial locations (see Figure A-2). The movement of the surge line was monitored with displacement transducers at 23 locations (see Figure A 3). Fifty-four thermocouples were installed, with either three, seven, or ten being placed at the nine axial locations. Each vertical location (locations 2, 4, and 13) contained three thermocouples placed symmetrically around the pipe. The horizontal locations (locations 1, 5, 7, 9, 11, and 12) contained seven

) thermocouples spaced from top to bottom down one side of the surge line.

Location 9 contained an additional three thermocouples evenly spaced along the other side of the surge line for a total of ten thermocouples at this axial location. The thermocouples were identified by axial location and relative position as follows: "12-T5" means the thermocouple at axial position 12, fifth from the top of the line.

Plant data has been recorded for four heatups, three cooldowns and during periods of power operation. The data was stored for the first heatup and cooldown (2/89) with a sample time of 20 seconds which was nearly the fastest update time possible for the Data Acquisition System. The subsequent heatups and cooldowns were primarily monitored with an update time of 60 seconds which was deemed acceptable based on the 20 second data. The large quantity of data (more than 180 Megabytes of disk space) was stored in ASCII format and transferred to the BWNS HP9000 Series 800 computer where the data was processed into functional information (plots, calculations, etc.).

O V A-4

l-Table _ A 1. . Oconee Stratification _ Test _ Signal List

~

SIGNAL SIGNAL DESCRIPTION NAME 1-TI-E- Surge Line Location 1, Thermocouple 1-E 1-T2-E Surge Line Location 1,; Thermocouple 2-E l-T3-E Surge Line Location 1,LThermocouple-3-E - i 1-T4-E Surgeiline Location 1, Thermocouple 4-E 1-TS E Surge Line Location 1,. Thermocouple 5-E 1-T6-E Surge Line Location 1, Thermocouple 6-E 1-T7-E Surge Line-Location 1, Thermocouple 7-E 10-L3 Y 3" LVDT, S/N-648.

10 SP X 5" Celesco String Pot,.S/N 25161  ;

10-SP-Y 5" Celesco: String Pot, S/N 2513_ '

11-L2-Y 2" LVDT, S/N:1996>

11-SP-X 5" Celesco String Pot, S/N-2520 -

11-T1-S Surge Line Location _11,; Thermocouple 1-S ll-T2-S Surge Line-Location 11, Thermocouple 2-S 11-T3-S Surge Line Location 'll, Thermocouple 3-S 11-T4-S Surge Line Location 11,. Thermocouple 4-S 11-T5 S Surge Line Location ll, ~ Thermocouple'5-S 11-T6-S Surge Line Location 11,: Thermocouple-6-S 11-T7-S Surge Line Location ll,- Thermocouple 7-S 12-L2-Y_ 2" LVDT, S/N-1995.

T1-S Surge Line Location 12,- Thermocouple 1-S 12-T2-S Surge Line_ Location;12, Thermocouple-2-S.

/' 12-T3 S Surge Line Location 12,- Thermocouple ~ 3-S 12-T4-S Surge Line Location 12, Thermocouple 4-S 12-TS S Surge Line Location 12, Thermocouple 15-S-12-T6-S Surge Line Location 12, Thermocouple.6 S 12-T7 S Surge-Line Location 12 Thermocouple 7-S 13-T1-W Surge Line Location 13,- Thermocouple 1-W 13-T2 NE Surge- Line Location- 13, Thermocouple NE-13-T3-SE Surge line Location 13, Thermocouple 3-SE 2-SP-X 5" Celesco String Pot,~ S/N:25181

=2-SP-Z - 5" Celesco' String Pot, S/N 2529.-

2-T1-N Surge Line Location 2, Thermocouple 1-N' 2 T2-SE Surge line Location 2,- Thermocouple 2;SE 2 I3-SW Surge-Line Location _2, Thermocouple;3-SW-4.. PIl52B : Loc 2-HPI NozzleLTemperature 152,1 Bottom-2HPI152T- Loc;2 HPI Nozzle Temperature 152,' Top 2HPI153B- Loc-2 HPI Nozzle Temperature 153, Bottom 2HPIl53T Loc 2 HPI Nozzle _ Temperature 153, Top 3-SP X- 5" Celesco String-Pot, S/N 2521-

'3-SP-Z 5" Celesco String Pot,1S/N 2522-A-5 1

Table A-1. (colit. )

SIGNAL SIGNAL DESCRIPTION NAME 3HPI1528 Loc 3 HPI Nozzle Temperature 152, Bottom 3HPIl52T Loc 3 HPI Nozzle Temperature 152, Top 3HPIl53B Loc 3 HPI Nozzle Temperature 153, Bottom 3HPI153T Loc 3 HPI Nozzle Temperature.153, Top 4-SP-X 5" Celesco String Pot, S/N 2527 4-SP-Z 5" Celesco String Pot, S/N 2525 4-T1-N Surge Line Location 4, Thermocouple 1-N 4-T2-SE Surge Line Location 4 Thermocouple 2-SE-4 T3-SW Surge Line Location 4 Thermocouple 3-SW 4HPI152 Loc 4 HPI Nozzl2 Temperature 152 4HPI153T Loc 4 HPI Nozzle Temperature 153, Top.

5-L2-Y. -2" LVDT, S/N 1977 5-SP-Z 5" Celesco String Pot, S/N-2514 5-T1-E Surge Line Location 5, Thermocouple 1-E 5-T2-E Surge Line Location 5, Thermocouple 2-E T3-E Surge Line Location 5, Thermocouple 3 E 5-T4-E Surge Line Location 5, Thermocouple 4-E 5-TS-E Surge Line Location 5 Thermocouple 5-E 5-T6-E Surge Line Location 5, Thermocouple 6-E 5-T7-E Surge Line Location 5, Thermocouple 7-E 6-L3-Y 3" LVDT, S/N 645 -

6-SP-Y 5" Celesco String Pot, S/N 2530 Os" 6-SP-Z 5" Celesco String Pot, S/N 2512 7-L2-Y 2" LVDT, S/N 1974 7-SP-Z 5" Celesco String Pot, S/N 2515 7-T1-E Surge Line Location 7, Thermocouple 1-E-7-T2-E Surge Line-Lucation 7, Thermocouple 2-E 7-T3-E Surge-Line-Location 7,-Thermocouple 3-E 7-T4-E Surge Line Location 7, Thermocouple 4-E T5-E- Surge Line Location 7, Thermocouple 5-E 7-T6-E Surge Line Location 7, Thermocouple 6-E 7-T7-E Surge Line Location 7, Thermocouple 7-E 8-L2-Y 2" LVDT, S/N 1988 8-SP-Z 5" Celesco String Pot, S/N 2531 9-L2-Y. 2" LVDT, S/N 1978 9 SP-Z 5" Celesco String Pot,.S/N 2519' 9-T1-E Surge Line Location 9, Thermocouple'l-E 9-T2-E- Surge Line Location 9, Thermocouple 2-E 9-T3-E Surge Line Location 9, Thermocouple 3-E 9-T4-E Surge Line Location 9 Thermocouple 4-E 9-T5-E Surge Line Location 9, Thermocouple 5-E 9-T6-E Surge Line Location 9, Thermocouple'6-E -

9-T7-E Surge. Line -Location 9, Thermocouple 7-E 9-T9-W Surge Line Location 9, Thermocouple 9-W A-6

_m,___ _ _ _ _ _ _ _.__.___.-_ ___m-.--- -

Table A-1. -(cont.)

SIGNAL SIGNA! DESCRIPTION NAME HPil52-1 .HPI Fazz 9 Temperature 152-1 HPil52-2 HP1 l'ozzle Tempei nure 152-2 HPil52-3 HPI hozzle Temperature-152-3 HPIl53-1 HPI hozzle Temperature.153-1 HPil53-2 HP1 tozzle Temperature 153 2 HPil53 3 .HPI N e.zle Temperature 153 HPil88T- HPI No2zle Temperature 188, Top-Lpzr Pressuiizer Level- .

LsgAop . Steam Menerator A Operate level LsgAsu Steam Uenerator A Startup Level LsgBop Steam tenerator B Operate Level LsgBsu- Steam G1nerator B Startup Level MT Trip Main Tutbine Trip Pms Miin Steim Pressure Presnr Narrow Range RC Pressure Preswr Wide Range Reactor. Coolant Pressure PsgA ' Steam Generator Pressure A i PsgB Steam Generator Pressure B

> Rate Ch2 Channoi 2. Rate of Change > 2% FP Ref L2 2"-LVDT, S/N 1976-5" Celesco String Pot, S/N 2528 A Ref-SP

-Reactor Power U RxPower TcldAwr Wide Range RC Loop A Cold Leg-Temperature TcidBwr Wide Range RC Loop B Cold-Leg Temperature Tcold avg Average Tcold Tdas Data Acquisition System Temperature Tfw Feedwater Temperature Thot avg Average That ThotA RC Loop A Hot Leg Temperature ThotB RC Loop B Hot leg Temperature Tmi Mirror Insulation Temperature Tpt String Pot Temperature TsgA Steam Generator A Temperature TsgB Steam Generator B Temperature WfwA Feedwater-Loop A Flow WfwB Feedwater Loop B Flow i Wrcs Reactor Coolant Total Flow l 1

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V Figure A-1. Oconee Data Acquisition Hardware Conf!gt: ration

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APPENDIX B B. Utilization of Oconee Data for BW Owners, Group. Program 1

l l- CONTENTS  ;

Page-

1. HPI Check Valve Test An Example Surge Line Transient . .-.-. . . .

B2

2. Operational Conclusions and Recommendations . . . .. -. . . . . . . . .- .

B 4' l

o b

t Y

O B1 .

9 s

Numerous observations pertaining to plant operations and the resulting effect on O surge line thermal stratification have been made based on the Oconee data. The effect of the HPl check valve tests on stratification discussed.in the first subsection provides an example surge line transient. This subsection provides ,

a more detailed view of stratification and the operations that cause it to occur.

The second subsection consolidates the conclusions drawn from the data.

1. HPI Check Valve Test - An Examole Surae Line Transient The most significant stratification cycles during the February 7, 1989 heatup 1 were due to High Pressure injection (HPI) check valve functional testing (2/10/89, 0100 to 0600). - A stratification cycle of 2SOF resulted from changes in pressurizer level at a time when the teoperature difference between the pressurizer and the reactor coolant system (RCS) was approximately 300F. Figure B 1 shows the relationship of key parameters during this test period and Figure B 2 is a diagram of the tested flow paths.

At Oconee, the HPI Borated Water Storage Tank (BWST) check valves (HP 10!.HP 102

- see Figure B 2) were tested for free movement to open by pumping from the BWST into the RCS (Oconee Functional Test PT/1/A/251/09 Enclosure 13.2). This test .

is required if it has not been performed in the last 90 days or if maintenance has been performed on either of the check valves. By monitoring the SWST level and the letdown Storage Tank (LOST) level while opening the shutoff valves (HP-24, HP 25) in the flow path be' tween the BWST and an RC cold leg, it could be determined if the suction check valves (HP-101, HP 102). were allowing flow.

M Ler opening HP 24, the operating HPl pump should take suction from the BWST.

At Oconee, the makeup control valve has been throttled open to establish greater than 30 gpm, or 150 gpni if following a refueling outage. After a- refueling outage,150 gpm of makeup flow is established to verify the check valves operate with a-full strokt f approximately 150 gpm is the maximum flow obtainable with the valves.in the full open position). If the BWST level decreases or the Letdown Storage Tank (LDST) level increases,. HP-101 was deterrined to be operating correctly. The test was repeated by opening HP-25 and testing for flow through HP 102 in the same manner.

The Oconee HPI BWST suction check valve testing had a large impact on surge line stratification, due to 1) large RCS inventory changes, and 2) the _ large

. O e.2 l

temperature difference between the pressurizer and the RCS at the time of the x testing. Because letdown flow is driven by the RC pressure, a maximum of 75 gpm was available at the time of .he HPl check valvo test. The pressurizer level was controlled at approximately 100 inches during this part of the heatup. Thus, an outsurge was initiated after each insurge to restore pressurizer levtl.

During these outsurges, the pressurizer temperature was between 340F and 41 W ,

while the RC temperature was only 100F (RC pumps had not been started). As a result of the outsurgos, stratification nearly as large as the pressurizer-RC temperature difference develops. Figure B 1 is a plot of key parameters for this time period with the detailed operational events annotated.

The 1A HPI pump was sthrted at n;40 and turned off 10 minut( 3 later. Inventory was added to the RC system as soon as tha pump was turned on. Because the pressurizer level was well above 100 inches at this i.ime, the makeup control valve was closed trying to control level to 100 inches. The makeup control valve has not been known to allow leakage, even with the large delta P across it in this early part of the heatup. Duke personnel suspect that the inventory addition had been due to seal injection throttle valve (HP-31) leakage. At this p time, the RC pressure (65 psig) was not sufficient to establish the proper seal d back pressure, the seal return throttle valve (HP 277) was adjusted for the l proper back pressure after an RC pressure of 295 psig was reached. This does not occur for another two hours. The pressurizer level fell when the HPI pump was turned off due to no inventory addition with a constant letdown of approximately 50 gpm.

The outsurges that occur between 0250 and 0520 were a result of the HPI BWST check valve testing. The insurges that begin at 0355 and 0415 were due to inventory addition during the HPI BWST suction valve test. Although not directly verified, it appears the same test was the reason for the other insurges occurring between 0250 and Ot?0. The evidence for this has been based on the similar time span for each insurge (5 to 10 minutes), the rate of pressurizer level increase, and the fact that these tests have often been redone for additional verification.

A portion of the RCS inventory changes up until 0530 were possibly due to leakage by the seal injection throttle valve (HP-31). This was due to the proper back pressure not being established until 0530. Adjustment of HP 277 to establish the

'O B-3

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proper seal back pressure occurred at 0531. The pressurizer level was relatively O. steady after that time.

The top to bottom temperature difference at the horizontal surge line locations at the time of the peak stratification (0449) was as follows:

Oconee Unit 1 Surge Line Location Approximate Top to Bottom .

Temperature Difference ___

j 5 275 l 7 190 f 200 )

11 280 l 12 250 l

The top thermocouples indicate approximately 400F for these locations. The bottom thermocouples at the two middle locations in the surge line (location 7 and location 9) were affected by the large outsurges but the bottom thermocouples at the other locations were not. This accounts for the lower peak stratification O t >#ree 14 e ioc tie #, 7 o 9.

The relationship between pressurizer level change and stratification can be seen in figure B 1. All decreases in pressurizer level cause the stratification to increase. However, stratification was more dependant on the rate of pressurizer level change than it was the total level change. This was evident at 0310 and 0330on2/10asshowninFigureB-1,

2. Operational Conclusions and Recommendations The surge line data taken at Oconee Unit I have confirmed that the design basis transients have not been adequate for the surge line structural analysis. The B&W Owners Group program has been established for the development of new surge line design basis transients in which the Oconee Unit I data collected played a key role in addressing surge line thermal stratification.

The Oconee data has indicated which plant operations are important in determining that the plants could be generically evaluated and for specifying the events to B4

-., ,.-3

be included in the design basis transients. The following conclusions in regard n v to operations have been drawn from the data:

1)

The RCS conditions in which the pressurizer to RC hot leg temperature difference was large have been determined to be of the most concern.

Power operations have been determined to generally be of less concern than heatup and cooldown events,

2) For outsurges, the surge line top temperatures increase with a smaller effect on the surga line botton. In addition, surge line bottom temperatures lag those of the top. Thus, quick changes in surge line top temperatures cause increased stratification before the bottom of the surge line responds. Insurges decrease stratification during most RCS conditions,

3) The first RC pump start results in a sharp outsurge due to reactor coolant mixing and subsequent contraction. Additional RC pump starts generally cause an insurge and pump stops cause an outsurge due to changes in loop pressure,

, 4) The operation of full pressurizer spray flushes the entire surge line and However, reduces stratification to a negligible magnitude. a I stratification cycle resulted because the surge line returned to its

( previous thermal condition after the spray flow was stopped, and

5) Operations which involve inventory additions might result in large stratification cycles. During the Oconee HPI check valve test, inventory was added to the system at a greater rate than it could be discharged by letdown. Af ter the inventory addition was stopped, restoration of pressurizer icvel resulted in an outsurge. The largest top to bottom temperature observed during the 2/89 heatup resulted from the HPI check valve test.

The October 2, 1989 heatup data (key parameters shown in Figure B 3) have been l

used to correlate thermal stratification to surge line boundary conditions. This computerized set of correlations developed by BWNS relates the transient surge line conditions to the imposed boundary conditions - (hot leg temperature, saturation temperature, and pressurizer level) as discussed in Section 4.2.

I k- B-5

d

'l Plant operation recommendations have been made based collectively upon the Oconee ,

-O surge line data and other available plant data. The first recommendation has .

been to implement a surge line pressure-temperature limit in order to limit the temperature difference between the pressurizer and the RC not leg. The second recommendation stammed from the HPI check valve test discussed earlier in this section. It has been recommended that surveillance tests that require large ,

makeup or letdown from the RC system be performed only when the RCS subcooled I margin is less than 200F in: order to limit the pressurizer to RC hot leg i temperature difference.

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APPENDIX C.

C. Supplementary Priping Information ,

J CONTENTS .

Page

1. Thermal Striping Literature . . . . . . . . . . . . . . - - . . . . . C.2 1.1. BWR Feedwater Nozzles Tests . . . . . . . . . . . . . . . . . . C4 1.2. LMFBR Tests . . . . . . . . . . . . . ... . . . . . . . . . ... C5 1.3. ANL Tests . . . . . . . . ... . . . . . . . . . . . . . . . . . . C6 1.4. HDR Tests ... . . . . . . . . . . . . . . . . .-. . .-. . . . . . C7

1.5. Comparison of Thermal Striping Conditions . . . .. . . . . . . . . C9 l- 2. Description of HDR Experiments ........_.......... C-10 l 2.1. Configuration . . . . . ... . . . . . . . . . . . . . ... . . . . -

C 11 2.2. Conditions . . . . . . . . . . . . . . . . .-. . .--... . . . . . C 11 2.3. Tests . . . . . . .-. . . . . . . . . . . . . . . . ... .-. . . . C-12 2.4. Th e rmo c o u pl e s . . . . . . . . . . . . . . . . . . . . . . . . . - . .

C 2.5. Data Acquisition Rate . . . . . . . . . . . . .-. .'. . . . . . . ..

C-13.

3. References ....:.....:................ .. C 14 l

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l~

O 1. Thermal Stricina Literature Thermal striping is the localized metal stress caused by repetitive fluctuations of the temperature at a fluid metal interface. The fluid temperature fluctuations are due to the interactions between forced flow aiid buoyancy. The buoyant forces tend to stratify the fluid, obtaining vertical segregation (in a horizontal flow component) by temperature and density. The fluid shear forces associated with forced convection and fluid viscosity, on the other hand, tend to mix the fluid. The combination of these effects can generate undulations of the fluid fluid interface, resulting in thermal striping. The existence and characteristics of these fluctuations depend primarily on the simultaneous occurrence of buoyant forces and fluid shear forces which are of comparable magnitude, The fluctuations are also responsive to the flow geometry; for example, they may be greatly amplified by the helical, secondary fluid motion induced by axial flow in a pipe bend. Finally, the fluctuations of the pipe surface temperature are directly responsive to the interactions between convective and conductive heat transfer at the fluid pipe interface.

The thermal-hydraulic characteristics of thermal striping have been examined in t2e rollowing four major areas of research:

1 BWR feedwater nozzle tests,

2. LMFBR tests (Westinghouse),

3. Argonne National Laboratory (ANL) tests, and

4. ProjectHDR(FRG) tests.

These programs are summarized below, and are described in more detail in the subsequent paragraphs.

The BWR feedwater nozzle studies (reference 1) were extensive, but the geometry of interest was quite unlike that of the pressurizer surge line. The BWR studies did demonstrate the ability to combine low temperature data with high-te.,perature data and with plant striping data, by suitably adjusting the low temperature results. The application of the BWR feedwater nozzle results illustrated the use of probability density - functions. Thermal fluctuations were analyzed to determine the frequency of occurrence of cycles having discrete ranges of C-2 s_ --

,.y., , . . , - - ._ __m.. . _ . , . . , . , - , , - - - ,

, ,,%, - m___ ,,.,,,..7

p amplitudes. These incremental-amplitude analyses were carried through the nozzle L/ stress analysis by introducing plant time at-conditions data. These BWR feedwater nozzle studies are described in detail in Section 1.

Woodward (reference 2) examined the fluid temperature fluctuations in transparent horizontal pipes. The amplitudes of the near wall fluid temperature fluctuations reached 60% of the imposed fluid temperature difference, with most of the cycles exhibiting amplitudes of 10 to 35%. The frequency of fluctuations ranged from 0.1 to 0.5 Hz. A film heat transfer coefficient was needed to determine the wall thermal fluctuations from those of the near wall fluid. Woodward referred to the studies of Fujimoto et al. Fujimoto et al (reference 3) also studied striping in a transparent horizontal pipe. A fluid density difference was imposed by adding calcium chloride to the warmer fluid stream. Thin squares of copper were used to measure wall striping. The striping amplitude was less than 10% of the imposed temperature difference (the temperature difference between the interacting hot and cold fluid streams). The film heat transfer coefficient was 1.25 to 7 times the Dittus-Boelter coefficient. The studies of Woodward, and of Fujimoto et al, are described in detail in Section 2.

Kasza et al at ANL (reference 4 14) have extensively tested thermal stratification and striping in transparent horizontal piping with bends in both the vertical and horizontal planes. Based on a limited amount of published '

power-spectral-density information, the higher amplitude fluctuations occurred at lower frequencies, 0.1 to 0.6 Hz, with amplitudes of 30 to 40% of the imposed temperature difference. The amplitude of fluctuations drops off sharply with increasing frequency. These ANL studies are outlined in Section 4.

Wolf et al (reference 15 27), in the TEMR test series of Project HDR, measured striping in metal ~, horizontal pipes at plant-typical temperatures. Typical striping frequencies were 0.1 to 10 Hz. The amplitudes of the wall temperature fluctuations were generally from 10 to 40% of the imposed fluid temperature dif-ference, with peak amplitudes from 25 to 50%. Examining a single test of the nine "PWR" tests, the maximum striping amplitude was approximately 30% and the frequency of occurrence of the larger fluctuations was approximately 0.2 Hz.

Wolf et al noted the interactions between convective and conductive effects, and hence the difficulty of extrapolating to a plant the results of tests performed using a transparent model. They also noted the insensitivity of temperatures C3 M

i

measured at the outside of a metal pipe to inside interactions. The TEMR tests 4

O are described later in Section 4.

Finally, the characteristics of thermal striping were addressed in a 1980 report by the NRC which summarized pipe cracking in PWRs. (referonce 28) The range of frequencies was 0.1 to 10 Hz. The reduction of amplitude due to film heat transfer was described, resulting '... in a peak metal temperature variation at the surface of roughly one fourth to one-half the water temperature variation."

1.1. BWR Feedwater Nozzles Tests

, The BWR feedwater nozzle configuration was examined in relation to observed feedwater line cracks. The thermal striping of this configuration has been obtained from two test facilities, Two Temperature and Moss Landing, as well as from plant measurements (reference 1). The Two Temperature Test Facility was limited to atmospheric pressure; hot and cold fluid temperatures of 160 and 70F were used. The Moss Landing Test Facility, on the other hand, achieved plant-typical temperatures. The results of the two test facilities were combined with plant data by adjusting _the Two-Temperature results to account for the changes of fluid density, viscosity, and thermal conductivity between the test' and reactor cenditions.

The composite data was processed to obtain the number of cycles having-discrete ranges of stress amplitudes. These amplitude ranges, or windows, were prescribed to be relatively small at the higher amplitudes, . 'p to 20% wide at the smallest amplitudes. The results of this analysis e a presented in tabular form. (reference 1) This data has been restated in terms of windows of equal amplitudes, 10%, and plotted in Figure C-1.- The frequency of occurrence decreases rapidly and regularly up to a stress amplitude of 50% of the maximum stress, and then more slowly at the higher amplitudes. Although most of the fluctuations had low amplitudes, approximately 1% of the metal temperature fluctuations obtained stress amplitudes approaching the maximum stress.

It is estimated that the observed amplitudes of the near-wall fluid temperature l

fluctuations were reduced by one-half to obtain the amplitudes of the wall temperature fluctuations and hence the wall stress amplitudes. The maximum

- stress of Figure C 1 thus corresponds to a wall temperature - fluctuation of l

approximately 50% of the imposed temperature difference, the temperature

'O C.,

. l

_ _ . _ . . ._ . . . . . . . . . - - . . . _ . . , _ , _ _ _ _ . _ _ . ~ _ ~. u.. _ ._ --w_.. -_.a.

i difference between the two fluid streams of unequal temperatures and densities.

O Most of the fluctuations had stress amplitudes less than 50% of the maximum-stress, which corresponds to wall temperature fluctuations less than 25% of the imposed fluid temperature difference.  ;

1.2. LMFBR Tests  ;

Woodward _(reference 2) investigated stratification _and striping _in a 1/5 scale model of an LMFBR at the Waltz Mills Test Facility. The model was plexiglass, therefore the hot and cold water temperatures were limited to 130F and 70F. Two lengths of horizontal piping, of 4" and 6.5" inside diameter, were examined. -The I

tests were conducted in the turbulent transition range, with Reynolds Numbers 3

(based on half pipe flow areas) of 2 x 10 to 8 x 103 . Dye and thermocouples

were used, the thermocouples were typically inserted 1/32' into the fluid.

The thickness of the interface region, over which the fluid temperature changed  ;

i from hot to cold, ranged from 0.6* to 2". The striping frequency was 0.1 to 0.5 Hz and the fluctuations were approximately sinusoidt.1. The amplitudes of the l temperature fluctuations (of the near wall fluid) approached 60% of the imposed temperature difference, and were most pronounced at low Richardson Numbers.

lO erobahiiitx of occurrence 4nformation was presented for enix three rances of amplitudes (or windows). This information has been converted to the fractional-occurrence for constant window widths of 10% amplitude, and plotted in Figure C. ,

2. Most of- the fluctuations had mid range ~ amplitudes, - 10 to 35%. The probability of occurrence dropped rapidly at the higher amplitudes, approaching zero et 60% amplitude.

A heat transfer coefficient was needed _to obtain wall temperature information from the near_wal_1 fluid temperature measurements. The - heat ' transfer coefficients determined by Fujimoto et al were referenced by Woodward. Fujimoto et al (reference 3) tested striping in a 14.2" horizontal pipe made of acrylite..

Calcium chloride was added to the warmer fluid stream to obtain plant typical density differences. (The fluid- temperatures were used simply to track the streams of differing densities.) Wall striping was measured on thin squares of copper. -The amplitude of the fluid temperature fluctuations was observed to

- decrease near the wall. The fluctuations at the interface between the fluids of

- differing densities evidenced frequencies of 0.3 to 3.0 Hz.. The amplitude of the -

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wall fluctuations was less than 10% of the imposed temperature difference. The O convective heat transfer coefficient was calculated to be from 1.25 to 7 times The that of the Dittus-Boelter correlation (for forced convection in tubes),

information obtained by Woodward and by Fujimoto et al was referenced in the evaluation of thermal stratification of the pressurizer surge lines of the South Texas Project power plants. .

J.3. ANL Tests Kasza et al have conducted extensive experimental studies of stratification and striping at ANL (references 414). These studies have generally used water flowing turbulently in transparent pipes of 6-in inside diameter. Combinations of horizontal and vertical piping lengths have been tested, including bends in the horizontal plane. Vertical lengths of piping were observed to interrupt stratification. Stratification in horizontal lengths began at a Richardson Number of approximately 0.05; flow stagnation and reversal occurred near a Richardson Number of 0.7. Kasza et al applied the buoyancy index of Jackson and Fewster, namely p T = Ri Re 0.625/ Pr l

Kasza et al observed a correlation between buoyant effects and this buoyancy index. The threshold of buoyant effects was found to correspond to a I on the order of 104; a T on the order of 10'2 or larger obtained strong buoyant offects.

The more recent investigations of Kasza et al (references 7 14) obtained some details of the thermal fluctuations. Whereas the bulk fluid temperature fluctuations were about 75% of the imposed temperature difference, the amplitude of the wall fluctuations was 30 a 40% of the imposed temperature difference.

On the basis of the limited published power-spectral-density information, most of the signal energy was concentrated below 1 Hz, peaking between 0.1 and 0.6 Hz, ar.d decaying approximately exponentially with increasing frequency. These maximum fluctuations were observed approximately one diameter downstream of a horizontal elbow. Kasta et al have obtained many additional, as yet unpublished, striping power spectral density distributions. These distributions confirm the trends of the published distribution.

(~h

' C6

1.4. HDR TestL Wolf et al have conducted extensive examinations of thermal mixing in the HDR project at Karlsruhe, FRG (references 15 27). The TEMB test series examined pressurized thermal shock using a large scale pressure vessel and various high-pressure injection configurations: the experimental results were compared to the predictions of many codes and correlations (references 15 22). Fluid temperature fluctuations were observed and recorded, but received little emphasis.

The TEMR test series concentrated on thermal stratification in horizontal feedwater lines (references 23 27). The test section was a 20 foot length of 15.6 in inside diameter metal pipe, extensively instrumented with Il ms thermo-couples. Cold water entered one end of the horizontal run through a bend from vertical upflow, the opposite end of the horizontal run was attached to a reservoir of hot fluid. The TEMR tests consisted of 3 subseries, 2 of which were labelled *BWR' and 'PWR." In the BWR tests, a plate with slit orifices was installed at the junction of the horizontal run with the reservoir, to simulate the holes of a typical E G feed sparger. The horizontal to reservoir junction was unobstructed in-the PWR tests. The third subseries of TEMR tests considered the buildup and decay of hot water pockets. The horizontal to-reservoir junction was blocked except for a horizontal slit at the bottom of the pipe cross section.

The PWR tests of the TEMR series are most relevant to the surge line con-figuration. The ranges of conditions of the 9 PWR tests are listed in Table C 1.

The average fluid temperature ranged from approximately 200 to 300F, and the imposed fluid temperature differences ranged from approximately 200 to 400F; the volumetric flow rates spanned 10 to 200 gpm. The Reynolds Numbers based on the flow area (rather than on a reduced flow area to account for stratification) were in the turbulent range. Kasza and Kuzay have employed a buoyancy index which is dependent on the Reynolds, Richardson, and Prandt1 Numbers. In an order of-ma0nitude sense, the threshold of buoyant effects occurs at an index of 10", and strong buoyant effects occur for an index of 10 and larger. Applying this index to the PWR test conditions, all the PWR tests of Wolf et al were in the strong buoyant range.

The interface between the fluid of unequal densities was characterized as being wavy,withtypicalfrequenciesbetween0.1and10Hz(reference 11). Within the

-C 7

.,,,..o,v.--.-e.,.-

mixing layer, the fluid temperature fluctuations were not damped near the wall q

V (reference 10). The fluid mixing did reduce the local maximum temperature d1tterence from the imposed temperature difference, however. The amplitude of the wall temperature variations, expressed as a fraction of the amplitude of the fluid temperature fluctuations, was stated in two contexts (reference 11). For all the BWR and PWR tests, the fractional amplitude was 10 to 40%, but the peak fractional amplitude was 25 to 50%.

Measurements from one of the PWR tests, Test 33.19, were presented (reference Test 33.19 was 11). The conditions of Test 33.19 are listed in Table C-1.

characterized by a relatively high flow rate and ratio of inertial to viscous forces (Re) and by a mid range temperature difference. The resulting ratio of buoyant to inertial forces (Ri) was low compared to that of the other PWR tests, as was the index of buoyant effects (1). The temperatures measured in the fluid, on the inside pipe metal surface, and on the outside pipe surface were presented.

(references 10,11) Examining these figures, the fluid temperature fluctuated with an amplitude which was almost equal to the imposed temperature difference, the difference between the temperatures of the hot and cold fluid streams. The temperature of the inside pipe wall fluctuated with an intermediate amplitude, but the temperature of the outside surface of the pipe evidenced no fluctuations.

The inside pipe surface temperature exhibited irregular fluctuations. The maximum amplitude of these fluctuations was approximately 30% of the imposed temperature difference; the larger amplitude fluctuations occurred at intervals of approximately 5 seconds. This interval corresponds to a frequency of occurrence (of relatively large fluctuations) of 0.2 Hz. This frequency of occurrence must be distinguished from the characteristics of the individual fluctuations. Because the larger-amplitude variations generally occurred within groups of fluctuations of much smaller amplitude, the frequency of all fluctuations was approximately 1 Hz. That is, the larger-amplitude fluctuations, which occurred at intervals of approximately 5 seconds, each persisted for only approximately 1 second. These characterizations were obtained by examining the figures presented for PWR Test 33.19.

C-8

The HDR experimentalist drew the following conclusions from the TEHR results:

O' (reference 10,11)

. The extrapolation of model data to a plant, using a transparent model, is made difficult by the complex interactions between convective and conductive phenomena,

. There is no simple, unique correspondence between the thermal response of the exterior of the pipe and that of the interior, and

. The region of fluctuating temperatures was extremely narrow (reference 26).

1.5. Comoarison of Thermal Strioina Conditions Figure C-3 provides an overview of thermal striping. Both dimensional and dimensionless axes are presented. The dimensional axes, flow rate versus temperature difference, apply specifically to the surge line geometry and conditions. The dimensionless axes, Reynolds number (Re) versus Grashof number (Gr), both correspond to the surge line quantities and provide a more general basis with which to assess -the thermal-hydraulic interactions. The Reynolds number indicates the ratio of inertial to viscous forces whereas the Grashof number provides a measure of the ratio of buoyant to viscous forces. The information presented in Figure C-3 is to be regarded in an order-of magnitude sense. For example, flow rates were converted to velocities using the whole pipe flow area, rather than reducing the area to accommodate stratification; and the surge line fluid properties were evaluated at 300F and slightly subcooled .they are relatively insensitive to pressure, but quite sensitive to temperature.

Regions of relatively weak and of relatively strong buoyant effects, compared to inertial effects, were estimated by evaluating the Richardson Number (Ri), and the buoyancy index (i) which has previously been described. The conditions of interest to surge line stratification and striping lie in the ' strong buoyant effects" range shown in Figure C-3. The range.of interest is further refined by i considering the surge line temperature difference (OT): there is no fatigue concern for DTs less than 90F, and the maximum DT is approximately-300F. The l

conditions of interest are thus approximately 10ll < Gr < 1012 and Re < 105 .

-(There is probably a lower-Re bound, below which the buoyant effects predominate to the extent that interface instabilities and thus striping are suppressed; this C-9

s

1 limit has not been quantified, except that Kasza et al have observed such a limit for the case of fluctuations downstream of a bend in the horizontal plane.)

The dimensionless axes of Figure C 3 provide a convenient basis on which to l

compare the several investigations of striping. These are the singly cross-hatched regions in the figure. The conditions of Woodward, and of Kasza et al,

! lie far below the range of Grashof numbers of interest. Both these striping investigations were conducted at atmospheric pressure, thus the reduced thermal expansion coefficient and DT, as well as the increased viscosity, resulted in relatively small Grashof Numbers. The conditions of Wolf et al, on the other hand, are just on the high Gr side of the conditions of interest, due only to their larger pipe diameter compared to that of the surge line.

Several data sets are not shown on the figure. The data of Fujimoto et al was obtained at atmospheric conditions, but with the inter fluid density difference enhanced toward that encountered at surge line conditions. The viscosity remained in the low temperature range, however, and the data evaluation of Fujimoto et al depended on a correspondence between mixing and diffusion within a thermal gradisnt and a concentration gradient. The EDF data-is also not shown.

This data, although unpublished, was apparently obtained at atmospheric pressure and correspondingly low Grashof numbers. Finally, the conditions of the BWR feedwater nozzle research are not shown because of the pronounced geometric dissimilarities between the nozzle and the surge line.

Notwithstanding their dissimilar Grashof number ranges, the published striping characteristics of the three investigations. plotted -in Figure C 3 are quite similar. But, it is uncertain whether the low Gr data of Woodward and of Kasza et al apply at the conditions of interest for the surge line. 'Certainly the visualizations available with the low-Gr tests provide valuable insight regarding i striping mechanisms, characteristics, and limiting regions, but these insights may apply only at the tested conditions, even if the Richardson number is preserved in the translation of conditions to those of interest. Hence, the work of Wolf et al thus seems singularly pertinent to surge line applications.-

i l 2. Descriotion of HDR Exoeriments l

l- The subject tests were conducted at Battelle Karlsruhe under the direction of Dr.

Lothar Wolf et al, in October '1986 (reference 29).- They were conducted in a C-10 T==5v mrytr up w e- -cm---e-y w -v v y -7 e--Aresqu a-f '?T-'-w-t---w=p-*e-ww pmii

. _ _ . , _ _ . _ ._ _ .. _ _ _ . _ . _ . - _ _ _ _ _ . _ . . . _ _ _ _ . - - -~_

i larger series of tests in horizontal piping with various configurations of the O flow area at the pipe to vessel interface.

2.1. Confiauration Stratification and striping were measured in an 18 foot, straight run of horizontal piping. Thepipeinsidediameterwas15.6in(397mm),itsthickness i was 0.634 in (16.1mm). The steel pipe (15NicuMoNb5, WB36) was insulated with glass wool. One end of the pipe was connected to a pressure vessel through a nozzle, and the opposite end to a gradual bend downward to a 21 foot vertical run (Figure C 4). The nozzle was unobstructed throughout the series of tests analyzed herein. The pressure vessel and horizontal pipe were initially hot and isothermal, cold water was introduced at the bottom of the vertical run during testing.

2.2. Conditions The ranges of imposed conditions were as follows:

Aversge Temperature (F): 187 to 323 Temperature Difference (F): 190 to 420 l O Flow Rate (gpm): 15 to 229 t

In terms of nondimensional variables, these ranges of conditions were:

Gr: 5 x 10**11 to 3.5 X 10**12 Re: 9.4 x 10**3 to 2.5 x 10**5 Ri: 2.7 x 10' to 1.1 x 10' T: 1.2 x 10** to 2.4 x 10' where 6 is the buoyancy index:

'6 Ri Re o.625 P r'

These dimensionless variables are global rather. than local; the imposed temperature difference rather than a local gradient was used, and velocity was calculated using the whole pipe flow area.

O C-ll 4

- - - - - - . - - , , , -m.- ,, y. v,r- , - - - ,..r,g - - . , --m-w-- , ,we-, - - - - . , , -c-w

?  !

, 2.3. Tests There were 7 tests conducted with an unobstructed pipe to vessel interface Tests 14 through 19 and Test 25. Test 25 consisted of 3 distinct phases, making a total of 9 sets of test conditions for analysis (Table C 2). The 3 phases of Test 25 have been separated, and reidentified as Tests 51 through 53,-for ease of analysis and identification (Table C 3). The imposed injection flow rate, fluid temperature difference, and average fluid temperature were varied among these 9 sets of test conditions. The warmer fluid was kept subcooled throughout each test.

Starti.ig with a hot and isothermal system, cold water injection was begun with a prescribed flow rate and temperature. Data recording was initiated as the cold to hot fluid interface ascended the vertical run toward the horizontal piping run. Time zero was set as this interface approached the. bend to the horizontal run, such that the effects of the imposed flow rate and temperature difference began to be observed just after time zero. Testing was generally continued until the fluid and metal temperatures approximately stabilized.

During this time, system pressure and the injection flow rate and temperature, O were to be held constant. The hot fluid was replenished by inherent (rather than imposed) backflow from the heated pressure vessel. The tests were terminated by

! interrupting the injection flow and stopping the data recording.

2.4. Thermocouples The horizontal pipe was instrumented with multiple scrain gauges and thermocouples. Displacement, pressure, and flow rate measurements were also performed. This investigation deals almost exclusively _with the temperature measurements, and usually with those at the instrument _ cross section located approximately two thirds of the way from the bend to the pressure vessel. This cross section was equipped with thermocouple triplets at- 30' increments from 0 to 180', at 15' increments from 105 to 120', and at 21/2' increments from 90 to 105' (Figure C-5). Each triplet consisted of a near wall fluid thermocouple, and thermocouples on the inside surface of the pipe and on the outside surface. The near wall fluid thermocouples were located 0.4 in (10mm) from the wall, compared to the pipe inside diameter of 15.6 in (397mm). The surface thermocouples were held to the wall using thin foils which covered the thermocouples almost to their O C 12 l

^

,-,,,,--.v-,,,-,.,,,l., . _ . _ , , . _ , _ _ _ , _ _

1-tips and which were welded to the wall. The diameter of the Chromel-Alumel '

thermocouple junction was 0.02 in (.05mm).

l 2.5. Data Acouisition Rate The data aralyzed herein was recorded at a frequency of 10 Hz. Examination of the temperature traces revealed that this frequency was completely adequate to document the temperature versus time fluctuationst multiple points were recorded l

for most of the fluctuations. The point density was sufficient to reconstruct the actual fluctuations with confidence. The predominant-striping frequency I obtained herein, approximately I Hz, supports the adequacy "of the 10 Hz '

measurements (but would itself have been biased had the predominant frequency ,

approached the frequency of data recording).

i d

O d'

C-13 9 e L--..

O

3. References

1. H. Watanabe,

• Boiling Water Reactor Feedwater Nozzle /Sparger Final Report," NEDO 21821-A (Februsry 80).

2. W.S. Woodward, " Fatigue of LMFBR Piping Due to-Flow Stratification," ASME Pressure Vessel and Piping Conference, Portland, Paper 83-PVP 59, CONF-830607 27 (June 83).

3. T. Fujimoto, K. Swada, K. Uragami, A. Tsuge, and K. Hanzawa, " Experimental Therma _1 Study of Striping at the Interface of Thermal Stratification,"

Hydraulict in Nuclear Technoloov, K.H. Sun et al (ed), ASME (81).

4. K.E. Kasza, J.P. Bobis, " Thermal-Transient Induced Pipe Stratification,"

6!ff, Li, 675 676 (80). ,

5. K.E. Kasza, J.P. Bobis, W.P. Lawrence, and T.M. Kuzay, " Heat Exchanger l

O ther , >eer eci <<rects:

ANL CT 81 31 (April 82).

oe 49e ee eerror ece cem eet <e8 >>.-

6. J.J. Oras and K.F. Kasza, " Thermal Transient - Induced Buoyant Flow i

Channeling in a Vertical Steam Generator-Tube Bundle," ANL-83-109 (Oct 83).

7. K.E. Kasza, T.M. Kuzay, and J.J. Oras, " Overview of Thermal-Buoyancy-Induced Phenomena in Reactor Plant Components," Proc 3rd Inti Conf -

QXf.91.d. pp. 187-194, (April 84).

8. K.E. Kasza, J.P. Bobis, W.P. Lawrence to J.C. Liljegren, " Thermal Transient Induced Pipe Flow Stratification; Phenomena and Correlations (Phase !!)," ANL-CT-81-19 (Feb 81).

9. K.E. Kasza and T.M. Kuzay " Thermal Transient Induced Pipe and Elbow Flow Stratification Phenomena and Correlations (Phase-III)," ANL 82 85 (Oct 82). .

C-14 y y r+-w-p-e ---,y- w y -. -

gw r-- v4r,--= , ,r ---w,,y

10. T.M. Kuzay and K.E. Kasza, " Thermal Oscillations Downstream of an Elbow in Stratified Pipe Flow, FBR Thermal Hydraulics, Trans ANS, 16, 794-796 ,

(June 84).

11. T.M. Kuzay and K.E. Kasza, " Experiments and Analysis of a Horizontal Pipe i Elbow in Stratified Pipe flow," Liouid Metal Thermal Hydraulics, pp 459-460.

12. T.M. Kuzay and K.E. Kasza, " Thermal Stricing Downstream of a Horizontal Elbow Under Thermally Stratified Flow Conditions," Joint Mta. ANS and AIF. 3 Washinaton, CONF-841105 10 (Nov. 84). ,

13. K.E. Kasza, J.J. Oras and R. Kolman, ' Measurement of Velocity Profiles in a Stratified Pipe Flow Recirculating Shear Zone using Laser Flow Visuali-zation" LLouid Metal Reactor Thermal Huraulics, pp. 458 460.

14. T.M. Kuzay and K.E. Kasza, " Resolution of Thermal Striping Issue. Down- ,

stream of a Horizontal Pipe Elbow in Stratified Pipe Flow," ANS Annual Meetina. Boston, C0HF 850610 4 (June 85), i

15. L. Wolf. K. Fischer, W. Hafner, W. Baumann, U. Schygulla, and K-H Scholl, O overview ef HDR tar 9e Scaie eTS Thermai Mixin9 Exper4ments and ^#alise, with 3 D Codes and Engineering Models," Trans 8th SMIRT,1, pp. 359 365 (85).

16. L. Wolf, U. Schygulla, W. Hafner, K. Fischer, and W. Baumann, *Results of Thermal Mixing Tests at HDR-Facility and Comparisons with Best Estimat.

and Simple Codes," Irans 8th SM1BI, E, BE.1-9 (85).

17. L. Wolf, U. Schygulla, F, Gorner, and G.E. Neubrech, " Thermal Mixing Processes and RPV Wall Loads for HPI-Emergency Core Cooling Experiments in the HDR-Pressure Vessel," HEQ, 21, pp. 337-362 (Oct. 86).

18. L. Wolf, U. Schygulla, W. Hafner, K. Fischer, W. Baumann, and T.G.

Theofanous, ' Application of Engineering and Multi-Dimensional, Finite Difference Codes to HDR Thermal Mixing Experiments TEMB," - proc 14th WRSRIM, NUREG/CP-0082 1, pp. 396 416 (Feb. 87).

l

19. L. Wolf, W. Hafner, U. Schyguila, W. Baumann, and W. Schne11 hammer,

" Experimental and Analytical Results for HDR-TEMB Thermal Mixing Tests for Different HPI Nozzle Geometrics," Trans 9th SMIRT, g, pp. 319 324 (87).

U C 15

20, U. Schygulla, E. Hansjosten, H.J. Bader, and K. Jansen, ' Assessment of Heat Transfer and Fluid Dynamics in Cold Leg and Downcomer of HDR TEMB Ev eriments," Itans 9th SMIRT, G, pp. 313-317 (87),

21. W. llafner, X. Fischer, and L. Wolf, " Computations of the NDR Thermal Mixing Experiments and Analysis of Mixing Phenomena," Ingns 9th SMIRT, E, pp. 301 312 (87).

22. L. Wolf, W. Hafncr, K. Fischer, U. Schygulla, and W. Baumann, " Application of Engineering and Multi-Dimensional Finite Difference Codes to HDR Thermal Mixing Experiments TEMB," HID,10A, pp.137-165 (June 88). ,

23. L. Wolf and U. Schygulla, " Experimental Results of HDR TEMR Thormal Stratification Test in Horizontal Feedwater Lines," Trans 9th SMIRT, D, pp. 361 366 (87).

24. L. Wolf, U. Schygulla, M. Geiss and E. Hansjosten, " Thermal Stratification Tests in Horizontal Feedwater Pipelines," Proc 15th WRSRl_M, NUREG/CP-0091, 1, pp. 437 464 (Feb 88).

L. Wolf, M. Geiss, W. Hafner, E. Hansjosten, and A. Talja, " Temperature 25.

O ne W i> Strain riecte t4oes oer4n9 the Tosa 18erm , str 14<4c t4en Te t-At HDR," 10th SMIRT, Anaheim, CA (Aug89).

26. L. Wolf, M. Geiss, E. Hansjoster,, and A. Tal,ja, " Evaluation Of the Thermal l- 2, Stratification' Tests at the HDR Facility," NURETH 4 Conference, Karlsruhe, FRG (Oct 89).

27. A. Talja and E. Hansjosten, "Results Of Thermal Stratific3 tion Tests In -a Horizontal Pipe Line At the HDR Facility." TRANS 10th SMIRY, E (89),

28. " Investigation and Evaidation of Cracking Incidents in ,oiping in Pressurized Water Reactors," NUREG 0691 (Sept 80).

l

29. U. Schygulla and L. Wolf, " Design Report - Thermsl Stratification Experiments in a Horizontal Piping Section,' HOR-Test group TEMR T33, ,

Battelle-FRG (November, 1986).

O C-16 ,

+r , -- -- , , , # y - yr, , , - , , , , , - , - = m t - , , e y

I t-Table C 1. HDR Test Series TEMR-PWR: ).anges i f Conditions and Conditions of >

f--

Test 33.19 (Ref. 24)

. The extreme conditions are listed for any of the 9 tests, rather theti for the tests having extreme combinations of conditions.

. The flow rates and dimensionless numbers use the flow area of the whole pipe; properties are evaluated at the average fluid temperature 4 obtains

. 1 is the buoyancy index used by Kasza and Kuzay, where f > 10 strong buoyancy effects.

Condition MlDimum Maximum Test 33.19 Fluid temperatures, F Hot 314 486 417 Cold 79 130 130 201 391 287 (Hot-Cold) 274 Average 198 290 Flow Rates Volumetric, gpm 10 200 200 Velocity, f t/s 0.016 0.34 0.34 a > io' 1.8 x to' O ae - ve/<-

Ri-gf6Td/vr (- 1/Fr') .

io' 5 x 10' 3 x 10' 5.5 x 10' Y = Ri Re-0.625j pga 0.025 84 0.026 l

5 l

C-17 O

k y-i

i Table C 2. Test Conditions (Entries Rounded) l q

QJ . , . . .

TE$i 14 15 16 17 18 19 51 52 53 315 313 413 417 417 486 485 484 l

Thot, F , b313 Initial  ! 61 74 77 63 77 85 65 66 117 Teold, F Haximum 62 98 123 65 111 148 66 116 162 l

Teold, F Temp 252 217 190 349 306 269 420 369 322 Diff.

l Thot-Max.

Teld, F Avg. Temp. 137 206 218 239 264 283 276 300 323 (Thot4 Mar.

.Teld)/2 ,

i Flow 0.77 5.4 10.1 0,72 5.4 10.2 0.88 6.0 10 Rate, Ibm /hx10

Pressure, 98 96 94 328 339 329 634 634 631 psia Gr x 10 0.50 0.59 0.61 1.5 1.8 1.9 2.8 3.2 3.5 j

Re x 10 O.95 7.4 15 1.2 10 _21 1.7 13 25 Ri x 10 2 56 1.1 0.27 100 1.8 0.45 91 1.8 0.55 Buoyancy 13 0.071 0.012 24 0.11 0.019 18 0.11 0.023 Index _

0- C-18 n r.. ~ - -- , . _ , , . , , . - ., ..--4 ,.

.- . - - ~ . - . . - . . . - . - . . - . . - . . . _ - .

~

Table C-3. Phases of Test 33.25 .

Times in seconds ,

The imposed boundary conditions were varied among the three-phases of Test 33.25, as shown in table C 2.

Ilmes Extracted from Test 33.25 Test Number' Starth Zero End Phase 51 -122.5 0 920 ,

1 ,

52 900 925 1195 2

3 53 1180 1200 1604.9 lO i

1 l-i

] Figure C 1, Frequency of Occurrence Versus Amplitude j

i 100 -

.. l Data beood on NEDO-21821 A (ref.1) adjusted for --

50 unworm ampinuo windows of 10*. -

1 20 -

,, _i j 10: -

cg- -

.5 -

o "- .

2 -

C _

1- -

~_ s.

0-0.5 I ' '

O 20 40 60' ' 80 - -100.

Amplitude (A), % of Maximum Amplitude f

O C-20

o. .- _ -. _ _____

4 O Figure C-2. Frequency of Occurrence Versus Amplitude of Near Wall Fluid Temperature Fluctuations Bar chart bened on Woodward (reference 2),

30 ,e' .,'ys ' .

" e'djusted for uniform ampledue windows of 10%

, ,' - s,

• - /  %

/-  %

/~ \

25 - 'e 5, l r i \

[ .\.

l'

)\

.20 -

I l \.-

.'s r

/

~

4> 4 r

\ r 15 --

/ ,

-\

.6 / e O :g.

LE

_ i-

. sg N.

\

10 -

- \'_

\.

N,

-g.

\

5 --

\-

-N_

\

N . g

's I l I 0

O -10 20- , 40 50 60 Amplitude (A), % of Maximum Amplitude -

4 C-21-.

._m m

~

3 o O O Figure C-3. Striping Conditions 6

- 10 Gr-(8voyant/ Viscous) forces Re -(Inertial / Viscous) Forces g g g 3

l - 100

, surgel1ne.

enditlons 5 --

of 10 '

interest p

3 4

ro m ,

gg ett i .

x . e,

/ st o O

u- f / / O

@s *

=

• y .

- 10 g ,,tt g

8 m

/ B 5/

3 10* ' -' E E

/

8. 8 g .

/

N $ 4 s . .:

EL -

8 5;

$b \

\\

E 9 10 Gr II 12 10 10 10 10 e n i Grashof Number (Gr) I 1 10 LOG

= n ,

Surgeline Temperature Difference, F

_ _ _. - _ ~ . . _ _ _ _

l-Figure C-4. Test loop Extracted from Figure 4.21. of reference 1.  !

Dimensions in m.

c' s

4 Pressure Vessel N, <

/g.

Y ,

Jong Measurement Cross Sectionst . ,, f OS CC / -

geu s

< i i

O  !

s, ,

.'y I

Y Cold Water Injection Site O

C-23

..- . . . . _ ._ _ _ _ . _ . _.. _ _ . _ _ _.._, ~ . . .. . ._. . _ _ _ .

~

4 Fioure C-4. Test loop O Extracted from Figure 4.2-1 of reference 1.c Dimensions- in m. *

(, , ,, t

^

Pressure Vessel s. <

'/#g

- c .< DOg a ~

s.

_MeasurementCrossSectionst.[4 ,/Og CC,g -

u-e ,

• 4~ j 4 -  :

, .. j

~,

. t

O  !

t l

i t

.t.

4

,.._,/ t

. . , g ' ,

.l.

N i( -i

,g/ -- Q -

-.4 w ;

J i

w

• , /

8

'-i

,y YI  !:

I i

I ' .

Cold Water. Injection Site 1 1 i

/ 8 O

C-24

c O r

,:.n ' .w aJ x ,,' s  ;

p lg f

w"t)4),\

Y

.. .A 90' e , ,,.pj. . . . . . . . . . . ... . . ,. __ . . ... ..... . . . . . . ,

5/

e i '  !*

o i

I,f;,

V;  ;

'[ l

,.r....'

W  ;

120' k,-  ; ,

135' e_; w i

j ws.m .

150' 180' sb.\

K/

Fioure C-5. Thermocouple Placement

/

At Measurement Cross Section CC '

Open circles denote fluid thermocouples. * .,\46' /j; Solid circles denote pipe inside surface and outside surface thennocouples. Dimensions in m. /

Extracted fmm Figures 4.2-5 and / 90' 4.2 6 of reference 1. y

• ~- '

~~~~~~~~~~~~~~~~~~~~*'*[*925' o .

o

• 95*

o 97.5' O

'/ 100' O

C-24

- _ . .